[
    {
        "image_filename": "designv11_63_0002935_iemdc47953.2021.9449597-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002935_iemdc47953.2021.9449597-Figure1-1.png",
        "caption": "Fig. 1. Three alternative rotor configurations. (a) Twin bearingless rotor. (b) AMB-SPM motor alternative 1. (c) AMB-SPM motor alternative 2.",
        "texts": [
            " The potential of oil-free magnetic bearingsuspended compressors (e.g., twin centrifugal compressor Turbocor TG310 from Danfoss, [12]) has been recognized in the IEA Newsletter, 2012 [13]. Turbocor has been optimized to work in related chiller applications (Danfoss 2018) [14]. This work presents a 160 kW bearingless motor applied to a residential heat pump application. A study of the competitive technology and a comparison with a more traditional AMB rotor is shown. The three rotor geometries to be compared are shown in Fig. 1. The first configuration has two bearingless units symmetrically on each side of the axial AMB disc. Two alternative configurations have separated radial magnetic bearings and an SPM electric motor unit. The difference between the two is the location of the axial AMB disc. The disc is either located outside the radial magnetic bearings (alternative 1) or inside them (alternative 2). II. ACTUATORS The actuators are designed in the FEM and optimized for performance and stress. For a comparison of the electromagnetic force, torque, and losses, the same mesh size constraints are imposed: shaft 4 mm, rotor yoke 2 mm, stator yoke 4 mm, stator teeth in SPM 2 mm, stator teeth in bearingless motor 4 mm, bandage and magnets 2 mm, and sliding mesh 360 points in three layers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001497_kem.861.113-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001497_kem.861.113-Figure6-1.png",
        "caption": "Fig. 6 Micro deformation mechanism of \u03b3 and \u03b3 'phases",
        "texts": [
            " [13] of Georgia Institute of technology in the United States established a crystal viscoplasticity (CVP) model of creep fatigue interaction for CMSX-8 single crystal superalloy. The model fully considered the micro deformation mechanism of \u03b3 and '\u03b3 phase. They pointed out that the small change of channel size has a significant effect on the plastic behavior of the alloy, which can predict the deformation behavior of the alloy under creep fatigue and thermo mechanical fatigue. Further research shows that the model can be used to quantitatively evaluate the effect of phase morphology changes on creep fatigue behavior (Figure 6). According to the creep curve of superalloy under constant load, Fu et al. [14,15] determined a set of parameters to predict the creep behavior of superalloy, and successfully predicted the creep life under other loads. Wang et al. [16] conducted thermomechanical fatigue experiments under different stress/temperature range, relaxation time and phase angle, studied the thermomechanical fatigue behavior, failure mechanism and life prediction method of the alloy. Through the comparative analysis of the fracture, it was found that the same phase thermomechanical fatigue mode was creep fatigue failure, and the different phase thermomechanical fatigue mode was oxidation fatigue failure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000068_edpe.2019.8883869-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000068_edpe.2019.8883869-Figure14-1.png",
        "caption": "FIGURE 14. Vibration pattern and vector diagram of radial natural modal of stator. (a) 2nd order. (b) 3rd order. (c) 4th order.",
        "texts": [
            " And the efficiency reaches up to its maximum, namely 90%, when the speed is rated speed (300r/min,110Hz). From charts above, the motor designed in this paper has a good performance. The Block Lanczos method of Ansys software is adopted to analyse the structural modal of motor. The modal steps of solution and extension are both 40. Not considering the influence of motor frame, 3-D modal FEM analysis of stator core and winding is conducted. Grid, vibration pattern and vector diagram of radial natural modal of stator are shown in Fig.13 and Fig.14. Natural frequencies corresponding to each vibration patterns are listed in Table VI. 8542 VOLUME 4, 2016 Four test points are located on the motor, and acceleration sensor is placed at each test point. The NO.1 and NO.2 test points are located at the feet of motor, and NO.3 and NO.4 points are located at the top of motor. The inverter switching frequency is 2kHz. Acceleration of each point is tested at rated speed and load. Tested accelerations of four points under different frequencies are listed in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure16-1.png",
        "caption": "Fig. 16 Boundary conditions of the box",
        "texts": [
            " The vibration response of the double-layer box was analyzed on the simulation platform. The material was set as structural steel with a density of 7850\u00a0kg m\u22123, a Poisson\u2019s ratio of 0.3, and an elastic modulus of 200\u00a0GPa. The meshed box is shown in Fig.\u00a015. Based on the actual structure of the double-layer box, fixed support constraints were set at the bolt holes of the outer box, and dynamic load was applied to the input bearing seat, star wheel carrier, and output bearing seat. The boundary conditions of the box are shown in Fig.\u00a016. The box body was subjected to load and vibration response was generated at different positions. Four points were selected at the output bearing seat, input bearing seat, and outer box foot as the evaluation points of the vibration response, which were named as evaluation point I-1\u20134, II-1\u20134, and III-1\u20134, respectively (Fig.\u00a017). Figures\u00a018, 19, and 20 show the time and frequency domain response diagrams for the different evaluation points. The peak acceleration and frequency of each vibrating body corresponding to Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002850_ssd52085.2021.9429493-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002850_ssd52085.2021.9429493-Figure1-1.png",
        "caption": "Fig. 1. Geometric structure of the kneed biped robot.",
        "texts": [
            " Using bifurcation diagrams, the trend of the variation of the walking parameters at different slope angles and also the variation of the slope angle under different values of the lower segment of the shank of the two legs, are analyzed in Section III. The periodicity of the passive gait of the kneed bipedal robot is also studied in this same section. Finally, we conclude this paper with a conclusion and some possible future works in Section IV. II. THE PASSIVE KNEED-GAIT BIPEDAL ROBOT This paper will present an analysis of the passive dynamic walking of the bipedal robot with knees in an attempt to mimic human walking and then in order to reveal further its complexity. Fig. 1 shows a geometric representation of the bipedal robot with knees descending an inclined walking surface with a slope parameter \u03d5. The kneed bipedal robot consists of two equal legs with knees, without ankles and with a punctual contact with the ground. As shown in Fig. 1, the length of each leg is L which is divided into two parts, L = ls + lt, with ls is the distance between the knee and the heel and lt is the distance between the knee and the hip. ms, mt and mh are the point masses concentrating at the shanks, thighs and the hip, respectively. Shanks and thighs are divided into two parts, ls = a1 + b1 and lt = a2 + b2 with a1 (resp. b1) is the distance between the sub-mass center 978-0-7381-4392-7/21/$31.00 \u00a92021 IEEE 179 20 21 1 8t h In te rn at io na l M ul ti- Co nf er en ce o n Sy st em s, S ig na ls & am p; D ev ic es (S SD ) | 9 78 -1 -6 65 4- 14 93 -7 /2 1/ $3 1",
            "1 10 9/ 3 20 21 1 8t h In te rn at io na l M ul tiCo nf er en ce o n Sy st em s, S ig na ls & am p; D ev ic es (S SD ) | 9 78 -1 -6 65 4- 14 93 -7 /2 1/ $3 1. 00 \u00a9 20 21 IE EE | D O I: 10 .1 10 9/ SS Authorized licensed use limited to: Cornell University Library. Downloaded on June 19,2021 at 01:15:13 UTC from IEEE Xplore. Restrictions apply. ms and the heel (resp. knee), and a2 (resp. b2) is the distance between the sub-mass center mt and the knee (resp. hip). Three other parameters depicted in Fig. 1 are: q1, q2 and q3, which are the angular positions, receptively, of the stance leg, the swing thigh and the swing shank. Two walking phases are necessary for the bipedal robot to complete its walking process as it is shown in Fig.2. These two phases are: The knee-free phase and the knee-locked phase. When striking, the knee will be locked and hence the kneed biped robot becomes similar to the compass-gait biped robot [25]. The values of the different parameters of the kneeling biped robot used in this work are shown in the Table I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002993_s12206-021-0629-6-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002993_s12206-021-0629-6-Figure2-1.png",
        "caption": "Fig. 2. Demonstration of lubrication angle \u03b1\u03c3 for worm drive.",
        "texts": [
            " The lubrication angle, induced normal curvature, and length of worm meshing line have different expressions in various coordinate systems to present a simple calculation. The lubrication angle and induced normal curvature will be calculated in S3i(O3-xiyizi), while the length of the worm meshing line will be calculated in S2i(O2-xiyizi). The lubrication effect could be reflected by lubrication angle well [21], whereas the lubrication angle \u03b1\u03c3 is defined by the intersection angle between the relative velocity of the worn drive and the tangent of meshing line (Fig. 2), v32 is the relative velocity for worm and worm wheel, and t-t indicates the tangent of the meshing line. High lubrication angle will lead to thick lubrication oil film and inevitably produce high transmission efficiency and long service life for the worm drive [22]. Therefore, high lubrication angle is beneficial to the worm drive\u2019s meshing performance. V32 is given as Eq. (10). According to Eqs. (5) and (9), the tangent of the meshing line for worm drive during the meshing could be expressed as: 3 32 3ib n\u03c9 \u2192\u2192 \u2192 = \u00d7 (14) Based on Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001026_ab9549-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001026_ab9549-Figure3-1.png",
        "caption": "Fig. 3. (a) Schematic showing parts of DEA and (b) fabricated DEA in unimorph configuration.",
        "texts": [
            " Composite films having varying CNT wt% were processed to optimize the CNT wt% in DEA. The sample nomenclature used according to CNT wt% in PDMS matrix is given in Table 2. Fig. 2. Processing steps followed to fabricate CNT/PDMS composite thin film. Table 2: Sample nomenclature according to CNT wt% in PDMS matrix. Sample ID Matrix CNT (wt.%) s1 PDMS 0 s2 PDMS 0.01 s3 PDMS 0.05 s4 PDMS 0.1 s5 PDMS 0.5 Unimorph DEA was fabricated by bonding a poly vinyl chloride (PVC) sheet on one side of the pre-stretched elastomer film as shown in Fig. 3. The extent of pre-stretch affects the actuation behaviour of DEA by increasing the dielectric breakdown strength and initial stiffness of elastomer layer. Based on the results reported in [25], a 30% of pre-stretch was used in all the experiments performed here. PVC sheet with 100 \u00b5m thickness was used to create a rectangular frame shown in Fig. 3(a) and pre-stretched elastomer film was glued to it. While carbon grease was used as electrodes, electrical connections on both sides of elastomeric film were made with copper tape. After removal of pre-stretching force, the elastomer film tried to regain its shape but the part of film bonded to PVC frame was not able to retract. Consequently, PVC frame and elastomer film formed a curved structure under new equilibrium condition. The complete assembly was clamped between rigid teflon grips to get unimorph configuration as shown in Fig. 3(b). 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d M an us rip t 5 As discussed above, an optimum concentration of CNT is required to enahnce the tip displacemnt of unimorph DEA. To this end, a detailed characterization of mechanical, electrical and actuation behaviour of CNT/PDMS composites processed with varying weight percent of CNTs is performed in this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002720_032043-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002720_032043-Figure1-1.png",
        "caption": "Fig. 1 Steering system model",
        "texts": [
            "1088/1742-6596/1885/3/032043 Steering system is a mechanism used to maintain or change the direction of the vehicle. When the vehicle is running, it ensures that there is a coordinated angle relationship between the wheels. The setting of the steering system in Trucksim is generally to select the steering shaft form, set the transmission ratio, and model the nonlinear characteristics of the system. The front suspension of the bus studied in this paper adopts a non-independent suspension, so the steering trapezoid is selected as an integral steering trapezoid. The model is shown in Fig.1. When a vehicle is braking, it is usually the driver who presses down the brake pedal to generate pedal force. The brake pressure enters the brake cylinder through the ABS system, and the braking torque is generated by fluid power to decelerate the vehicle. Fig.2 shows the working process of the braking system. According to the relevant test parameters and standards for passenger vehicles, the maximum braking torque of the front axle is selected in Trucksim as 7.5 kN\ua78fm; The maximum braking torque of the rear axle is 10 kN\ua78fm, and the ABS braking system is selected in the braking module"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000508_robio49542.2019.8961531-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000508_robio49542.2019.8961531-Figure2-1.png",
        "caption": "Fig. 2. The object O is represented by a coordinate frame (x, y, z) that fixed it, and a grasp pose g is represented by the grasp center position p, the grasp width w, and the grasp angle \u03d5 around z-axis.",
        "texts": [
            " [5] directly evaluated the grasp quality of consecutive pixels in the image space and simultaneously regressed the grasps angle and width with a single neural network. This architecture has a very fast processing speed, but each pixel has one fixed grasp angle and width. Also, it has lower accuracy than the grasp evaluation methods. In order to overcome the challenging problem of localizing robot grasp configurations directly from the point cloud, Hongzhuo Liang et al. [22] introduced their convolutional neural network called PointNetGPD which can directly process the 3D point cloud. We use a specific parallel gripper to grasp objects and, as shown in Figure 2, we define a grasp state in the world coordinate as g = (p, \u03d5, w) \u2208 R3 \u00d7 S1 \u00d7R1 (1) where p = (x, y, z) \u2208 R3 is the position of grasp center point in cartesian space, w \u2208 R1 is the maximal open width of gripper. As the grippers are always perpendicular towards the work plane, we define \u03d5 \u2208 S1 as the rotation around the z axis. To transform a grasp pose from cartesian space to image space, we have the depth image defined as I = (o,To,Kc,Tc) \u2208 RH\u00d7W + (2) where H ,W are the height and width, subscript o represents the property and geometric shape of the object to be grasped, c represents the parameters of the camera, To and Tc are the position matrix of the object and the camera in the world coordinate respectively, Kc is the camera\u2019s intrinsic parameter matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure2-1.png",
        "caption": "Fig. 2 Three-dimensional model of the double-layer box",
        "texts": [
            " According to the rotor\u2013stator coupling relationship between the encased differential gear system and the double-layer box, the vibration energy of the output bearing, input bearing, and star gear bearings is transferred to the output bearing seat, input bearing seat, and star wheel carrier of the double-layer box, respectively. Therefore, the output bearing seat, input bearing seat, and star wheel carrier are the excitation sources of the vibration system of the double-layer box. The three-dimensional model of the double-layer box, shown in Fig.\u00a02, is mainly composed of the inner and outer 1 3 boxes. The outer box provides support to the inner one through two sets of vibration isolators. The connection diagram of the different parts is shown in Fig.\u00a03. The main components are the outer box, the output bearing seat, the inner bracket 1, the inner bracket 2, the star wheel carrier, the input bearing seat, and the vibration isolator. The outer box is supported at its base; thus, it is the vibration receiving structure of the vibration system of the double-layer box"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000608_s00170-020-05023-4-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000608_s00170-020-05023-4-Figure3-1.png",
        "caption": "Fig. 3 (a) Cross section of adjacent cladding passes. (b) Cross section of arbitrary cladding layer in overlapping direction",
        "texts": [
            " In the following subsections, two deposition strategies in both directions are expounded. In overlapping direction, the height difference of each cladding layer will be realized by dynamically adjusting the overlapping rate. In the existing flat-top overlapping methods [17], the overlapping rate usually remains constant during forming process. In this section, an oblique-top overlapping method is presented to generate diverse height difference between adjacent cladding passes by discretely adjusting the overlapping rate. As shown in Fig. 3(a), to ensure the surface flatness, the overlapping area A1 should be equal to A2, meaning that the trapezoidal area AEFHG = (ABEAD1 + ABEAD2)/2. Meanwhile, the area of Bead 1 also satisfies ABEAD \u00bc 4H2 \u00feW2 8H 2 arcsin 4HW 4H2 \u00feW2 \u2212 W W2\u22124H2 16H \u00f02\u00de where W and H are the width and height of a cladding pass. Thus, the center distance of adjacent beads is defined as C \u00bc ABEAD1 \u00fe ABEAD2\u00f0 \u00de= h1 \u00fe h2\u00f0 \u00de \u00f03\u00de Subsequently, the maximum and minimum height hOt and hOb of each layer in overlapping direction are solved by hOt \u00bc HOT=N hOb \u00bc HOB=N \u00f04\u00de where HOT and HOB are the lowest and highest height of the whole part in overlapping direction. As shown in Fig. 3(b), through recursive computation, the height hk + 1 of kth pass can be deduced by the former height hk and the sum of center distances C1~Ck. Moreover, hk + 1 can be also solved by Eq. (2) and Eq. (3). Therefore, the number of cladding pass can be recursively calculated by hk\u00fe1 \u00bc ABEADk \u00fe ABEAD k\u00fe1\u00f0 \u00de =Ck\u2212hk hk\u00fe1 \u00bc \u2212 hOt\u2212hOb L0\u2212W \u2211C1::k \u00fe hOb \u2211n k\u00bc1Ck \u00bc L0\u2212W 8>< >: \u00f05\u00de To ensure a constant powder defocusing value, a \u201cbottom-up\u201d method is proposed to generate the scanning path of jth layer. Based on the NURBS surface model Mtop, the height HPj at arbitrary point Pj on the jth slicing surface will be generated and satisfies the following: \u0394hj \u00bc j HST\u2212HSB\u00f0 \u00de=N \u0394hPj \u00bc f LP \u0394hj \u00fe HSB=N HPj \u00bc HPMtop\u2212\u2211N j \u0394hPj;HPj\u2208 jHSB=N ; jHST=N\u00bd 8< : \u00f06\u00de where\u0394hj is the maximum height difference of jth layer,\u0394hPj is the thickness at Pj of jth layer, fLP is the linear interpolation function between\u0394hj and\u0394hPj, and HPMtop is the height of P on surface Mtop"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002829_5.0052213-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002829_5.0052213-Figure1-1.png",
        "caption": "FIG. 1. Streamline plot of the velocity field in the low (top) and high (bottom) curvature regimes, in response to a force localized on the spherical membrane. On the left, the flow field is shown in a h;/ chart while on the right, the flow field is wrapped on a spherical membrane. Location of the force is marked in red. Note the creation of two vortical defects around the force. In the top row (low curvature regime), we show the flow field for a point force localized at h \u00bc 1:5. In the bottom row (high curvature regime) for a point force localized at h\u00bc 0, the vortices migrate to the equatorial regions.",
        "texts": [
            " There are two distinct regimes R > k (low curvature) or R < k (high curvature). (The ratio k=R is often quoted as the Boussinesq number in the surfactant dynamics literature.) In the low curvature regime, the velocity field exhibits a dipole-like structure around the point of application of the force. The dipole has a topological index\u00fe 2 which agrees with the Euler characteristic of the sphere. As the curvature is increased, the dipole structure breaks into two\u00fe 1 vortices which migrate away to diametrically opposite points (Fig. 1). These features were predicted first in Refs. 52 and 53 and generalized to lipid bilayers with slip velocity in Ref. 57. We observe that our real-space Green\u2019s function (Appendix A) also reproduces these effects. This provides a consistency check of our summation procedure explained in Appendix A. Similarly, for a force dipole, one expects the flow field to be characterized by four vortical defects surrounding a saddle of negative index at the core of the dipole. There must exist an additional saddle of negative index such that the net index is\u00fe 2, the Euler characteristic of the sphere"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003025_tee.23411-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003025_tee.23411-Figure1-1.png",
        "caption": "Fig. 1. Equivalent circuit model of IPMSM",
        "texts": [
            " With the FSMPC, the proposed strategy performs well in tracking the MEA trajectory and improving efficiency and extending the operating region of the IPMSM. 2. Equivalent Circuit Model of IPMSM The IPMSM model, including the voltage equation, the electric torque function, and the power loss function, is the foundation for optimizing efficiency and achieving the FSMPC. In general, the equivalent circuit model is often used to analyze the IPMSM considering the iron loss [20]. The equivalent circuit model of the IPMSM is shown in Fig. 1. The continuous-time voltage equations of the IPMSM in the rotating dq-frame can be given as follows. { ud = Rs id + Ld diod dt \u2212 \u03c9eLq ioq uq = Rs iq + Lq dioq dt + \u03c9e(\u03d5f + Ld iod) (1) where \u23a7\u23a8 \u23a9 id = iod \u2212 \u03c9e Lq ioq Rc iq = ioq + \u03c9e (\u03d5f +Ld iod) Rc (2) The stator resistance (Rs ) can be measured directly by a digital multimeter. Ld and Lq are the dq-axis inductances. The equivalent iron loss resistance (Rc) is defined as (3). The parameters (Ld , Lq , K e , and K h ) can be estimated with the finite-element analysis (FEA) results and the experiments [21]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure9-1.png",
        "caption": "Fig. 9. Instantaneous volume between rotor and stator and the projection of areas.",
        "texts": [
            " The output of the module is provided in TXT format and can be used for stand-alone considerations of the motor geometry and as input to the fluid dynamics & force module. The module is divided into two submodules: the gear geometric module and the commutator geometric module. The gear geometric module evaluates several features required for evaluation of fluid-dynamics features such as instantaneous TSV volume, port areas, hydraulic diameters. A sample TSV (colored in green) of the rotor-stator set is presented in Fig. 9. The module evaluates the volume using GSL libraries. Apart from the evaluation of geometric features, the gear geometric module also evaluates the projection areas of each TSV along both X and Y axes, Lx b; Ly b (Fig. 9) respectively, and the points of application of these forces. The projection areas are essential to evaluate the fluid pressure force applied to the rotor. The points of application of these forces are essential to evaluate the torque applied to the rotor because of the fluid pressure forces. All these quantities, essential to evaluate both the fluid and the force features are provided as a lookup table with respect to the rotation angle of the orbit motor. The commutator geometric module evaluates only the porting characteristics realized by the commutator with respect to the windows manifold (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002250_s12666-020-02152-y-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002250_s12666-020-02152-y-Figure7-1.png",
        "caption": "Fig. 7 Semi-automatic design of methoding (feeding and gating systems)",
        "texts": [
            " It empowered even novice engineers (who lack foundry experience) and senior engineers (who may have limited computer skills), to effectively and efficiently use it without elaborate training. Their feedback helped in improving successive versions of the software. Over the next 10 years, the software was also employed to train several hundred foundry engineers in casting design and simulation through continuing education programs organized at IIT, Bombay. The tenth version of the software was re-developed with an entirely new user interface, additional features, more automation and expanded alloy database (Fig. 7). It was named AutoCAST-X, and launched at the World Foundry Congress at Chennai in 2008. Five years later, an advanced version of the software (AutoCAST-X1) incorporated the FLOW? module mentioned earlier, which was launched at the Indian Foundry Congress in Ahmedabad. Till date, it has been implemented at about 200 organizations. Most of them are foundries; others include OEMs, R&D institutes and consultants providing simulation services. The endusers have reported good matching of simulated and observed results, as well as ease of optimizing the methoding, leading to overall improvement in casting quality, yield and energy utilization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure13-1.png",
        "caption": "Fig. 13 Three-dimensional graph",
        "texts": [
            " When the mechanism is in configuration 3, the spring force at kinematic joint D makes the components 3 and 4 to be relatively static. Based on the principle of augmented Assur groups, the planar doublefolded metamorphic mechanism is divided into a fixedaxis rotating active part and an augmented Assur group RRRP (Fig.\u00a012b and c). 1 3 According to the geometric and physical parameters of the planar double-folded metamorphic mechanism in Table\u00a02, a three-dimensional model is established in SolidWorks, as shown in Fig.\u00a013. The initial location of component 4 is 227.73\u00a0mm, and the initial position of the component 1 is \u03c0\u00a0rad. Assuming that the active part 1 rotates at a constant speed of 6 r/min (motion period is 10\u00a0s), the dynamic simulation is carried out in SolidWorks virtual prototype environment, and the relationship between the driving torque and time of the planar double-folded metamorphic mechanism is obtained. When the planar double-folded metamorphic mechanism is in configurations 1 and 3, the mechanism under force constraint can be regarded as consisting of an active part and an augmented Assur group RRRP (as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure2-1.png",
        "caption": "Fig. 2. 2D model of Permanent Magnet Synchronous Machine.",
        "texts": [
            "), the PMSM can be shared into its symmetry parts. This means, instead of doing the simulation of the complete motor, only one of its symmetric components is simulated and then the complete motor behaviour can be extrapolated. Fig. 1 shows the RMxprt model of the studied machine. When the RMxprt model was analysed, it was exported to the two-dimensional and three-dimensional model using the ANSYS Maxwell Software. In this case, the machine has been divided into four equal parts and only one part is represented on 2D and 3D model as shown in Fig. 2 and Fig. 3. The geometrical of 2D and 3D model are depicted in Fig. 2 and Fig. 3. 682 Authorized licensed use limited to: Raytheon Technologies. Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. During the FE simulation of permanent magnet synchronous machine, the meshing is essential process which is done to descretize the geometry developed into small number of parts called cells. A quarter geometry of the PMSM is subdivided into triangular elements (1000 elements). As obvious from Fig. 4 and Fig. 5 both models give relatively the detailed view of the mesh quality in the 2D and 3D model of the PMSM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002223_s0263574720001368-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002223_s0263574720001368-Figure2-1.png",
        "caption": "Fig. 2. The local coordinate system of the ith limb.",
        "texts": [
            " (1) B0 RA0 = Rot(z, \u03c6z)Rot(y, \u03c6y)Rot(x, \u03c6x) (1) where \u03c6x , \u03c6y , and \u03c6z are the Euler angles of rotation of the moving platform on the fixed x-axis, y-axis, and z-axis, respectively. The linear velocity, angular velocity, and angular acceleration of the moving platform can be expressed as follows:23, 24 v = q\u03070 (2a) \u03c9 = [ \u03c6\u0307x \u03c6\u0307y \u03c6\u0307z ]T (2b) \u03c9\u0307 = [ \u03c6\u0308x \u03c6\u0308y \u03c6\u0308z ]T (2c) The generalized position, velocity, and acceleration of the moving platform can be defined as follows: x = [ q0 \u03c6x \u03c6y \u03c6z ]T (3a) x\u0307 = [ v \u03c9 ]T (3b) x\u0308 = [ v\u0307 \u03c9\u0308 ]T (3c) To describe the orientation of each limb, the coordinate system Bi \u2212 xi yi zi is established at point Bi as illustrated in Fig. 2. The rotation matrix B0 RBi , which describes the relationship between the coordinate system Bi\u2212xi yi zi and the coordinate system B0 \u2212 xyz, can be written as follows: B0 RBi = Rot(z, \u03c6i )Rot(y\u2032, \u03d5i ) (4) The position equation associated with the ith limb can be written as follows: q0w0 + ai = bi + qi wi i = 1, 2, 3 (5) where q0, qi , w0, wi , ai , and bi represent the length of central limb, the length of the ith limb, the unit vector along the central limb, the vector along the ith limb, the vector A0 Ai , and the vector B0 Bi , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure2-1.png",
        "caption": "Fig. 2. Exploded view of the TL240 unit: 1. ports block, 2. manifold on the shaft side, 3. floating spline joint, 4. rotor-roller set, 5. windows manifold, 6. commutator. The arrows highlight the flow passages inside the TL 240 unit.",
        "texts": [
            " Section 3 presents the simulation model developed in this study for the evaluation of the fluid dynamic and force features, followed by the details of test set-up (Section 4). Section 5 discusses the major results and validation of the simulation model with experiments, in terms of mean and instantaneous torque, followed by conclusions. The main information on the Parker Hannifin\u2019s TL 240 unit taken as reference in this work is listed in Table 1 [13]. The exploded view of the reference orbit motor and the six major components of it are shown in Fig. 2. The ports block, 1, has provision to provide the inlet flow to the machine and to collect the outlet flow. The components 2,4,5 and 6 are stacked together by studs. The slots for the studs are larger than the cross-section of the studs, so that a flow passage is provided from the ports block, 1, to the commutator, 6. The floating spline joint converts the orbiting motion of the rotor to a rotation of the shaft, thus acting as a universal joint. The rotor-roller set comprises the core of the orbit motor, converting the pressure energy to mechanical energy",
            " The windows manifold and the commutator, together, perform the porting action neces- sary for the orbit motor to displace the working fluid. The windows manifold provides connection to the inlet/outlet environment with each of the inter-teeth volumes (also referred as Tooth Space Volumes, TSVs), whose definition and evaluation will be detailed in Section 3. The unit TL240 is bidirectional but, for the sake of clarity, the description as well as the study performed in this paper refer to the flow direction shown in red (input)/blue (output) arrows of Fig. 2. The inlet flow, represented in red in Fig. 2, enters from one of the ports in ports block, 1, and it passes through the slots before being supplied to the outside of the commutator, 6. The outlet flow, represented in blue, is provided to the inside of the commutator by the windows manifold, 5. The inside of the commutator is directly connected to the through hole of the spline joint, 3, through a slot in the latter. The through hole of the spline joint is connected to the other port of the ports block. Therefore, the commutator separates the inlet and outlet environments",
            " Three views of the windows manifold are presented in Fig. 3. In these views, the diametrically outer slots represent the passages through which the studs pass through. The flow enters or leaves through the inner slots presented in Fig. 3(a), as dictated by the commutator. Fig. 3(c) represents the view of windowsmanifold when viewed from the rotor side. Each one of the inner slots represented in Fig. 3(c) is connected to each of the TSV in the rotor-roller set. The commutator and rotor are connected by the floating spline joint (Fig. 2) and thus they move synchronously. For the continuous operation of the orbit motor, high pressure fluid is provided at a different angle to the TSV, as represented by the cross-sectional view of Fig. 3(b). Fig. 4 represents the windows under consideration along with the commutator (colored in gray) over an orbit rotation. The commutator performs the orbiting action, subjecting each of the windows in the windows manifold, 5 (Fig. 2), to the inside or outside of itself. From Fig. 4, it can be observed that the windows are exposed to the inside of the commutator or the outside which correspond to the outlet (colored in blue) and inlet environments (colored in red) respectively. For a certain duration of angle, the window area is completely enveloped by the commutator thus providing zero cross-porting. The structure of the developed simulation model is presented in Fig. 5. The four modules of the simulation model are: the geometric pre-processor module, the mechanical pre-processor module, the gap module and the fluid dynamics & force module",
            " The projection areas are essential to evaluate the fluid pressure force applied to the rotor. The points of application of these forces are essential to evaluate the torque applied to the rotor because of the fluid pressure forces. All these quantities, essential to evaluate both the fluid and the force features are provided as a lookup table with respect to the rotation angle of the orbit motor. The commutator geometric module evaluates only the porting characteristics realized by the commutator with respect to the windows manifold (Fig. 2). As explained earlier, the commutator separates the high-pressure and the low-pressure environments thus subjecting each of the windows to high and low pressures sequentially while the orbit motion is being performed (Fig. 4). The commutator geometric module evaluates the instantaneous area of communication and hydraulic diameter of each of the windows using GSL libraries. Fig. 10 represents part of the output generated by the commutator geometric module. For one of the windows, the instantaneous port communication area to the inlet and outlet environments is presented for one orbit revolution of the rotor",
            " Upon evaluation of these quantities at all the possible contact points, the values are stored in a lookup table for subsequent use. Such table can be used to interpolate for contact force values, arm-lengths and normal angles at a given position of the rotor. This approach can be applied to any machine, but the values of Dxmax and Dxdiff are not known a priori and a choice of them can only be justified after obtaining the maximum deviation of the machine observed, after the implementation of fluid dynamics & force module. Fig. 16 represents an exploded view of the motor with the components numbered in accordance to Fig. 2. The gap module evaluates the leakages across the lubricating interfaces (arrows in Fig. 16) of the orbit motor. In particular, there are two moving elements (ME), the rotor and commutator, in conjunction with stationary elements (SE). Lubricating interfaces exist on either side of the orbiting rotor (4). Also, a lubricating interface exists between the commutator (6) and windows manifold (5). A sealing element is present between the commutator and the casing, thus effectively sealing the gap. The gap height of these lubricating interfaces is evaluated from the lengths of different components and by assuming equal gap height distribution on either side of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002348_s12206-021-0133-z-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002348_s12206-021-0133-z-Figure10-1.png",
        "caption": "Fig. 10. Fixed platform and a pattern attached to an extensible limb.",
        "texts": [
            " It should be noted that parameter ir is also computed by the procedure. However, this geometrical parameter is not relevant for calibration purposes. On the other hand, point iS generates a spherical surface F iV , with center in iA . These points are located at the inter- section of the two axes of revolution of its corresponding chain, and further define the dimensions of the fixed platform, as shown in Fig. 1. Therefore, in order to determine the coordinates of points iA relative to the reference frame of the camera, an experiment is mounted as shown in Fig. 10. It is noteworthy that, for this procedure, it is not necessary to activate the actuator. Points C iA are determined in the same way that the coordinates of c iS . As shown in Fig. 2, the mobile platform is geometrically defined as the polyhedron whose vertices are the centers of rotation of the spherical joints. Following the methodology previously described, it is possible to locate these points since they correspond to the centers of the vertex spaces shown in Fig. 11. As mentioned before, these points are fitted to the corresponding sphere using the least squares method, as it is graphically illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001663_icuas48674.2020.9214003-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001663_icuas48674.2020.9214003-Figure1-1.png",
        "caption": "Fig. 1. Position of each rotor of the quadrotor.",
        "texts": [
            " The equations that represent a generic UAV [12] can be described as follows: p\u0307np\u0307e p\u0307d = c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8  uv w  (1) u\u0307v\u0307 w\u0307  = r v \u2212 q wpw \u2212 r u q u\u2212 p v + g \u2212s\u03b8c\u03b8 s\u03c6 c\u03b8 c\u03c6 \u2212 1 m 00 f  (2) \u03c6\u0307\u03b8\u0307 \u03c8\u0307  = 1 s\u03c6 s\u03b8/c\u03b8 \u2212c\u03c6 s\u03b8/c\u03b8 0 c\u03c6 \u2212s\u03c6 0 s\u03c6/c\u03b8 c\u03c6/c\u03b8  pq r  (3) p\u0307q\u0307 r\u0307  = I\u22121 pq r \u00d7 I pq r + \u03c4\u03c6\u03c4\u03b8 \u03c4\u03c8  (4) where cx = cos(x) and sx = sin(x). m, g and I denote the total mass of the UAV, the acceleration of gravity and the inertia matrix, respectively. \u03c4 = [ \u03c4\u03c6, \u03c4\u03b8, \u03c4\u03c8 ]T denotes the total external moments applied and is obtained from (7). By using the UAV\u2019s geometry represented in Fig. 1, the total thrust f can be computed as f1 + f2 + f3 + f4. The torque values \u03c41 and \u03c42 are computed according to the axes\u2019 directions in Fig. 2. This corresponds to (5)-(7), where d denotes the UAV diameter and mi denotes the torque generated by each propeller. Aerodynamic forces are considered perturbations of the system, since they have small influence compared to the ones produced by the motors and the mass of the UAV. \u03c4\u03c61\u03c4\u03b81 \u03c4\u03c81  = d 2  \u2212 \u221a 2 2 \u221a 2 2 \u221a 2 2 \u2212 \u221a 2 2\u221a 2 2 \u221a 2 2 \u2212 \u221a 2 2 \u2212 \u221a 2 2 0 0 0 0   f1 f2 f3 f4  (5) \u03c4\u03c62\u03c4\u03b82 \u03c4\u03c82  =  0 0 \u2212m1 +m2 \u2212m3 +m4  (6) \u03c4\u03c6\u03c4\u03b8 \u03c4\u03c8  = \u03c4\u03c61\u03c4\u03b81 \u03c4\u03c81 + \u03c4\u03c62\u03c4\u03b82 \u03c4\u03c82  (7) UAV\u2019s mass was found by using a scale and is m = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure11-1.png",
        "caption": "Fig. 11. Geometry of the bearing for the FE models.",
        "texts": [
            " Mart\u0301In et al. Tribology International 161 (2021) 107098 making geometrical simplifications based on symmetry planes, thus reducing the computational cost of the simulations. In this last aspect, only one sector with two roller halves was modelled for the axial load case, taking advantage of cyclic symmetry; for radial and tilting moment load cases, only half bearing was modelled, since the geometry, the boundary conditions and the loads are symmetric. The geometry of the studied bearings is shown in Fig. 11, with the geometrical parameters of Table 2 [4]. Rings were dimensioned following the standard geometry proposed in [18], but dependent on the housing H rather than on the rolling element diameter Dw [12]. Fig. 12 shows the efficient model developed in this work. The diagonal lines represent the MATRIX27 or COMBIN39 elements, which are connected to the rigid shells in the contact surfaces (in red) by means of rigid beam elements. It can be observed that these models do not require refined meshes which, together with the fact that rollers and wires are not considered, significantly reduces the DoF and the analysis cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000114_elecsym.2019.8901529-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000114_elecsym.2019.8901529-Figure3-1.png",
        "caption": "Fig. 3 Type of Bearing Faults",
        "texts": [
            " Type 6201 RS is used because based on the type of induction motor used, bearing 6201 RS is a compatible type to be installed in the Moswell induction motor model AQUA DB-125. Following are the Bearing 6201 RS specifications in Table I: Table I is the Bearing 6201 RS specification which is the type of Bearing used in this research. Referring to the research conducted by [4] [8] [15] [16] Bearings must be conditioned under normal conditions and fault. So that in this research Bearings are conditioned in the category of three fault, namely outer damage (Outer race), damage to the ball (Ball), and damage to the inside (Inner race). More details like in Fig. 3: Fig. 3 explains where in Bearings the outer condition is damaged in the outer layer. Bearing Inner fault condition was damaged in the inner layer and the last Bearing was damaged Ball condition was damaged in the Ball Bearing section. Bearing fault detection system in this study uses an MMA 7361 accelerometer sensor. The position of the accelerometer sensor is very influential on the results of vibration readings performed. The sensor must be placed in the place closest to the Bearing so that the vibration results obtained are vibrations capable of representing the influence of the condition of the Bearing itself"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001683_icuas48674.2020.9213979-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001683_icuas48674.2020.9213979-Figure2-1.png",
        "caption": "Fig. 2. Top view of the three dimensional path follower based on the carrot chasing algorithm without taking into account the orientation error.",
        "texts": [
            " Setting these parameters properly is key to get a good performance, depending on the desired 4D trajectory. A much more dense list of waypoints is required to apply the trajectory following method efficiently. Hence, it uses the trajectory generator to get a discrete curve \u03b3(\u03bb) from the ordered list of waypoints W . Then, continuously, it receives the UAS pose p(t) and the actual time t to generate the velocity commands v(t), based on the method described below. The proposed trajectory following method is based on the carrot chasing algorithm, see Fig 2. The method runs as follows, first the \u03bbp argument is obtained using the expression 1, which minimizes the distance from the UAS position to the trajectory. The fixed look ahead distance d is added and the virtual target pose in the trajectory is achieved as pt(t) = \u03b3(\u03bbp(t) + d). (4) To fix the time error, the cruising speed is calculated for the time t as vc = d \u03c4(\u03bbp(t) + d)\u2212 t . (5) The last step is to calculate the velocity command, based on the cruising speed, as v(t) = vc pt(t)\u2212 p(t) |pt(t)\u2212 p(t)| (6) to reach the target virtual pose"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure7-1.png",
        "caption": "Fig. 7. Density model domain which will be subjected to topology optimization.",
        "texts": [
            " (3) In cartesian coordinates, i, j \u2208 {x, y, z} and u is the material deflection. Equation (1) is subjected to boundary conditions, being either S = 0 (free boundary) or u = 0 (fixed boundary). These respective conditions are indicated on the geometry shown in Figure 6. Since the application requires some boundaries to be maintained in place, i.e. the motor enclosure and the back end of the non-conductive holder, meant to press on the battery pack, the optimization is performed on the domain shown in Figure 7. In addition to the deformable material, we have modelled the battery pressure by a rigid domain subjected to a prescribed displacement of 0.5 mm while being free to rotate. Solving eq. (3) above gives the classical solutions to mechanical problems without any optimization. Using the SIMP method, the stiffness D0 is replaced by a spatially dependent stiffness D(r) = \u03c1(r)pD0, (4) where p > 1 is an adjustable parameter (selected to be 3, as commonly done [25]). The stiffness D0 for anisotropic materials consists only two independent material parameters, where the Poisson\u2019s ratio, \u03bd, has been set to 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000879_j.promfg.2020.04.145-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000879_j.promfg.2020.04.145-Figure3-1.png",
        "caption": "Fig. 3. Temperature distribution on the substrate during welding at 65s.",
        "texts": [
            " In fixation strategy 4, all four corners of the substrate are constrained. In fixation strategy 5 the transversal edges are clamped along with the four corners (in the z-direction) to improve the possible longitudinal distortions of strategy 2, while an additional middle node on the lower side of the substrate is constrained in x and y-direction. The distortion in the welding is the result of the difference in the thermal shrinkage of the material due to an inhomogeneous temperature distribution [16] as shown in Fig. 3. The temperature distribution on the substrate during welding clarifies that the welding region on the substrate is at a much higher temperature than the area surrounding it. On cooling, the region which is at higher temperature shrinks more than the other region that is at a lower temperature. On the other hand, the rapid increase and decrease in the temperature also develops residual stresses on the substrate around the beads [17]. Hence minimizing the distortion as well as residual stresses is an important design criterion of the welding simulations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure13.21-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure13.21-1.png",
        "caption": "Fig. 13.21 Small-scale vehicles equipped with BeagleBone Black boards",
        "texts": [
            " It is similar to the one at KTH Stockholm [2], but uses a significantly cheaper motion capture system and is dedicated to autonomous ground vehicles only. The hardware and software components of the setup are described subsequently. Two different kinds of small-scale vehicles have been acquired: trucks for scenarios such as parking with trailers or platooning, and passenger cars to test standard maneuvers such as ACC or overtaking, which are described subsequently. Trucks Three trucks3 in scale 1:14 are available on the testbed, where two of them are shown in Fig. 13.21a. The vehicles have been equipped with a BeagleBone Black,4 which is a low-cost single-board computer. A WiFi module5 has been installed in order to receive the actual position in the room or to communicate with other vehicles. The trucks are actuated by a motor and servos for gear changing and steering via PWM signals. The steering angle is within the range |\u03b4truck| \u2264 17\u25e6 and the maximum allowable speed of the trucks has been set to 1.5m/s due to indoor testing and thus limited space for the testbed, but speeds may also be larger in different settings. 3http://www.tamiya.de/de/produkte/rcmodelltrucks.htm. 4http://beagleboard.org/black. 5http://www.tp-link.com/at/products/details/cat-9_TL-WR802N.html. Passenger Cars Three cars6 in scale 1:10 as shown in Fig. 13.21b have been built up, which are also equipped with a BeagleBone Black and a WiFi module. However, the actuation upon acquisition differed from the trucks: the speed of the included motors strongly depended on the state of charge of the battery and did not yield accurate results. Hence, brushless DC (BLDC) motors with Hall sensors7 and motor speed controllers8 have been used to achieve better performance at low speeds. In addition, the transmission gears have been replaced,which also improves driving at low speeds",
            "at/de/reely-strassenmodel-audi-rs6-brushed-110-rc-modellauto-elektrostrassenmodell-allradantrieb-rtr-24-ghz-238002.html. 7http://www.amainhobbies.com/lrp-vector-x20-brushless-motor-6.5-lrp50674/p226984. 8http://vedder.se/2015/01/vesc-open-source-esc. 9http://www.logitech.com/de-at/product/c930e-webcam. 10http://gopro.com all links accessed: 2018-04-09. 11http://www.optoma.de/projectorproduct/x320ust. purpose, the positions of moving objects are estimated via webcams and AprilTags, which are mounted on the vehicle as shown in Fig. 13.21. The position tracking code is based on the AprilTags C++ Library,12 which provides fast and robust 3D position estimation and the tracking algorithm is insensitive to bad light conditions. The camera intrinsics have been calibrated using AprilCal [25], which is necessary only once for each webcam. To cover the testbed area used for the presented experiments, four webcams mounted on the ceiling have been used for this position estimation. Note that for driving on a planar environment, it is sufficient to estimate a 2D position; hence, for the calibration of the camera extrinsics, a planar calibration target based onAprilTags has been used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure8-1.png",
        "caption": "Fig. 8 The general structure of the PHYSIOTHERABOT.",
        "texts": [
            " The central processing unit performs the following tasks as common to all types of exercise: - Algorithm selection - Send patient parameters that are stored in the database to the algorithms - Send the data from the sensors to the rule base and database - Send information from the rule base and the database to the controllers The developed robotic system has three DOF and can perform flexionextension for the knee and flexion-extension and abduction-adduction movements for the hip. Servo motors are used to actuate the mechanism and force sensors are used for force measurement. It can be adjusted according to the limb size and is suitable for both legs. The general structure of the robot manipulator is given in Fig. 8. The knee link (Link 2) of the PHYSIOTHERABOT was designed according to the parallelogram principle. The robot manipulator\u2019s motors are placed on the base to reduce the effects of motor weights on the robot manipulator dynamics. 413Impedance control applications in therapeutic exercise robots 414 Erhan Akdogan and Mehmet Emin Aktan The electronics hardware block diagram of the PHYSIOTHERABOT is given in Fig. 9. System hardware consists of servo motors, reductors, motor drivers, data acquisition cards, encoders, force sensors, limit switches, and emergency stop buttons"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001541_ccc50068.2020.9189251-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001541_ccc50068.2020.9189251-Figure3-1.png",
        "caption": "Figure 3: Forklift motion analysis",
        "texts": [
            " then affects the angle of the wheel through the hydraulic power steering system. Following the forklift dynamics, the coordinates (xUWB, yUWB) and heading angle ( ) are adjusted as a consequence, which are measured by the positioning and orientation system (UWB and AHRS). As a foundation for controller synthesis, the three degree of freedom dynamic model for the forklift is developed with rear wheel steering system, modified from the existing models proposed by [16]. In particular, as illustrated in Figure 3, the dynamic model is listed in equations (1) to (4). \u0307 = \u2217 ( ) \u2212 \u2217 (1) \u0307 + = \u2212 ( \u2212 ) + + (2) \u0307 = ( \u2212 ) + \u2212 (3) \u0307 = (4) where is the yaw rate, is the vehicle sideslip angle defined as the longitudinal axis of the vehicle and the orientation of vehicle velocity vector, and V is the vehicle velocity. is the angle of the wheel, a, b are the lengths from the G (Center of Gravity) to the front and rear axles respectively, m is the mass of the vehicle. The forces acting on the vehicle are the front lateral force , the rear lateral force , and the front longitudinal force (driving forces)",
            " The NMP behavior is the first important phenomenon observed in the relation from heading angle to the lateral movement. As shown in Figure 4, the forklift is headed in 90 direction from 0 to 21 s, and the y coordinate stayed constant since there is no movement in this direction. Then the forklift switched to 160 degree after 21s. An interesting observation is that the y coordinate increased in the initial stage, and gradually decreases after 22s. This is the wellknown NMP phenomenon. To understand this phenomenon physically, the following analysis was carried out as illustrated in Figure 3. The derivative of the lateral error of the measurement location (M) to the target trajectory (red dash line in Figure 3) can be calculated as shown in equation (11). is assumed to be 0. is target heading from target trajectory. e\u0307d= ( \u2212 ) + \u0307 ( \u2212 ) (11) where L is the length from the coordinate\u2019s measurement point M to the forklift\u2019s center of gravity. When \u2212 is minor, then ( \u2212 ) can be approximated as \u2212 , ( \u2212 ) can be approximated as 1 . This reduces (11) into: e\u0307d= ( \u2212 ) + \u0307 (12) Since is the control input, and ed is the controlled output, the transfer function for (12) can be approximated as: ( ) ( ) = + (13) Obviously, when L is negative (M is behind G) there is a right-half plan zero, leading to NMP behavior",
            " For forklift, it is common to move along straight line(s). For a straight line starting from A (xa,ya) to B (xb,yb), the lateral distance (ed) from the forklift (x,y) to the line (A to B) can be calculatea as[18], ed= yb-ya *( -xa)+ y-ya *(xa-xb) (ya-yb)2+(xa-yb)2 (14) The ideal heading direction along the target line can be obtained as, =arctan (xb-xa) (yb-ya) (15) Note that (x,y) is a reconstructed position estimated based on the measured position (xUWB, yUWB) and measured heading direction as observed from Figure 3. = + \u2217 cos ( 2 \u2212 ) (16) = + \u2217 ( 2 \u2212 ) (17) For the outer-loop controller, the lateral error for the reconstructed point to the target line follows the dynamic in equation (18), e\u0307d= sin( + \u2212 ) (18) where is the velocity at the reconstructed point (approximated by V), is the angle between the forklift velocity at the reconstructed point to the heading direction (approximated by ) .This rewrite (18) into, e\u0307d= sin( + \u2212 ) + ( \u2212 )sin ( + \u2212 ) + \u2206 = ( + \u2212 ) + (19) where \u2206 is the error caused by the discrepency between and , between and "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000015_aim.2019.8868337-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000015_aim.2019.8868337-Figure4-1.png",
        "caption": "Figure 4. wheelbase change strategy for overcoming obstacle",
        "texts": [
            "3  (b) The relationship between the wheelbase and the height of obstacle that can be overcome for front axle 1.1 1.3 1.5 1.7 1.9 2.1 0.2 0.4 0.6 0.8 1 1( )l m 0.5  0.4  0.3  h /r (c) The relationship between the wheelbase and the height of obstacle that can be overcome for middle axle Based on above analysis, dividing the crossing process into three parts: front axle wheels crossing obstacle, middle axle wheels crossing obstacle and rear axle wheels crossing obstacle, then wheelbase adjusting strategy (shown in Figure 4) for crossing obstacles can be derived: moving the middle axle forward to reduce the wheelbase l1 before the front axle wheels crosses obstacle (Figure 4a), after the front axle cross the obstacle (Figure 4b), wheelbase l1 remains unchanged and the middle axle wheels continues to overcome obstacles (Figure 4c). After the middle axle complete overcoming obstacle (Figure 4d), moving the middle axle wheels back to increase the wheelbase l1 (Figure 4e), so the crossing obstacle ability of the rear axle wheels can be enhanced. Finally, the vehicle finishes overcoming the obstacle (Figure 4f). Adopting such wheelbase adjusting strategy, the ability of overcoming obstacle of each wheel will be significantly improved, thereby improving the passing ability of the vehicle. The advantages of wheeled vehicles compares with tracked vehicles are their high mobility, low energy consumption and low noise. The drawbacks are that the passing ability and load carrying capacity are not as strong as those of tracked vehicles. In order to make up for the defects and retain the advantages of wheeled vehicle, a six-wheeled vehicle with variable wheelbase is proposed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure2-1.png",
        "caption": "Fig. 2. Temperature distribution",
        "texts": [
            " Model reference Properties Components Name: Ti-6Al-4 V alloy SolidBody (<Base alloy> <Fillet1 > ) Model type: Linear Elastic Isotropic Default failure criterion: Max von Mises Stress Yield strength: 8.27371e + 08 N/m^2 Tensile strength: 1.05e + 09 N/m^2 Elastic modulus: 1.048e + 11 N/m^2 Poisson\u2019s ratio: 0.31 Mass density: 4,428.78 kg/m^3 Shear modulus: 4.10238e + 10 N/m^2 Thermal expansion coefficient: 9e-06/Kelvin and 36,521 elements respectively. The model material properties are stated in Table 1. The temperature distribution on the base alloy at 293 K is shown in Fig. 2. While comparison of mesh quality plots is indicated in Figs. 3 and 4. This was done at room temperature in order to pre-heat the base metal and check the stresses via modelling. Pre-heat treatment of base alloy is good because it allows it to absorb the laser beam uniformly and avoid residual stresses that can tamper with the microstructures with larger grain sizes and influence the surface quality of the additive manufactured component. In addendum, component load resistance can be tampered with when the residual stresses are larger"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000481_sensors43011.2019.8956820-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000481_sensors43011.2019.8956820-Figure4-1.png",
        "caption": "Fig. 4. Schematics of QMACS setup.",
        "texts": [
            " The effluent of the plasma jet is imaged onto the entrance slit. The spectra are recorded with an Andor Istar Electron multiplying charge-coupled device. Three maps of the OH(A-X) emission intensity integrated from 305 to 313 nm are recorded, one for dry feed gas and humid shielding gas, one for dry shielding gas and humid feed gas, and one for dry feed gas and dry shielding gas. B. IR Quantum Cascade Laser Absorption Spectroscopy The ozone density has been investigated by quantum cascade laser (QCL) spectroscopy in the infrared region [10], [11]. Fig. 4 shows the complete diagnostic setup with the kinpen and the gas shielding device. The measurement system is based on the Q-MACS system developed by neoplas control GmbH, which has been optimized for operation at atmospheric conditions [12]. The midinfrared light source is a pulsed QCL emitting in the spectral range from 1024 to 1030 cm\u22121. The radiation is controlled by temperature adjustment. The QCL operates in single mode. The system can be tuned within a small range of 0.4 cm\u22121 by an increasing of the operating current",
            " It is then focused onto a spherical mirror with a focal length of 1 m and guided through a 60-cm length multipass white cell. After 28 reflections within the cell, resulting in a total absorption length inside the multipass cell of 16.80 m, the intensity is measured by a fast mercury cadmium telluride detector. A detailed description of the measurement procedure can be found in [10] and [12]. Briefly, the measurements were performed at atmospheric pressure in a glass tube with openings at each end for the mirrors, where the midinfrared reflection of the beam passes without disturbance. As shown in Fig. 4, the plasma jet is positioned in the middle of the multipass cell, without the laser passing through the plasma itself. The plasma source feed gas is blown directly into the glass chamber. The species distribution is assumed to be homogeneous. Fig. 5 shows a simulated and a recorded spectrum of the ozone absorption at standard pressure and temperature conditions. The spectral region of interest, corresponding to the maximum tuning range of the QCL source, ranges from 1026.8 to 1027.2 cm\u22121 (shaded area in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure8.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure8.2-1.png",
        "caption": "Fig. 8.2 Schematic diagram of a flow over an instrumented surface",
        "texts": [
            " 8.8. Notation: For any real p \u2265 1, L p(J ) denotes the set of Lebesgue-measurable functions x : J \u2282 R \u2192 Rwith finite norm \u2016x\u2016L p(J ) = (\u222b J |x(s)|p ds) 1 p , and L\u221e(J ) denotes the set of Lebesgue-measurable functions with finite norm \u2016x\u2016L\u221e(J ) = ess supt\u2208J |x(t)|. In this section, we reproduce the example given in [5] on how a bilinear delayed differential equation can be obtained as an input\u2013output model for a controlled flow system. Consider a flow over a surface as depicted in the scheme in Fig. 8.2. The actuator is an on\u2013off air blower and the sensor is a hot film [5]. To describe the behavior of the flow, we consider the Burgers\u2019 equation \u2202v(t, z) \u2202t + v(t, z) \u2202v(t, z) \u2202z = \u03bd \u22022v(t, z) \u2202z2 , (8.1) where v : R2 \u2192 R is the flow velocity field, z \u2208 R is the spatial coordinate, and \u03bd \u2208 R+ is the kinematic viscosity. Recall that (8.1) can be considered as a unidimensional approximation to the Navier\u2013Stokes equations for an incompressible flow [4]. Assume that, for some F \u2208 R+, z \u2208 [0, F]where z = 0 is the position of the actuator and z = F is the position of the sensor",
            " With these relations and by defining x(t) = v(t, F), u(t) = v(t, 0), (8.2) can be rewritten as x\u0307(t + \u03c2)=\u2212 1 F u(t \u2212 \u03c2)x(t) + 1 F x(t + \u03c2)u(t) + 4\u03bd F2 [ x(t) \u2212 2x(t + \u03c2) + u(t) ] , or equivalently, x\u0307(t) = \u2212 1 F x(t \u2212 \u03c2)u(t \u2212 2\u03c2) + 1 F x(t)u(t \u2212 \u03c2) + 4\u03bd F2 [x(t \u2212 \u03c2) \u2212 2x(t) + u(t \u2212 \u03c2)] , where \u03c2 = F/(2c). Hence, in [5\u20137], the authors propose the more general equation x\u0307(t) = N1\u2211 i=1 ai x(t \u2212 \u03c4i ) + N3\u2211 k=1 \u239b \u239d N2\u2211 j=1 c j,k x(t \u2212 \u03c4\u0304 j ) + bk \u239e \u23a0 u(t \u2212 \u03c2k) , (8.3) as the input\u2013output model for the separated flow control system shown in Fig. 8.2. The measured output is x and the control input is u. Observe that this approximating model still recovers two main features of the original flow model (8.1): first, it is nonlinear; and second, it is infinite dimensional. It is important to mention that for several experimental settings, this input\u2013output model replicates with good accuracy the input\u2013output behavior of the physical system, see [5] for more details. We consider again the controlled system depicted in Fig. 8.2 and its input\u2013output model given by (8.3). The two main concerns in the control design process of such a system are These two tasks must be done taking into account the following physical restrictions of the system. (a) The actuators are on\u2013off air blowers; thus, the image of the input signal u is restricted to the set {0, 1}. (b) A sensormeasurement x(t) = 0 indicates that theflow is in a separated condition, and a measurement x(t) > 0 describes a reduction of the flow separation. Thus, the higher the sensor measurement, the less the flow separation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002396_0142331220987532-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002396_0142331220987532-Figure1-1.png",
        "caption": "Figure 1. Representation of the different winding faults of 3-ph stator (model 3/3).",
        "texts": [
            " Model (3/3) three-phase to highlight stator faults Various forms of faults can have stator windings. However, four common stator faults are mainly present: turn-to - turn fault, short Phase-to - Phase fault, short Phase-to-Earth fault, and open circuit coil fault. These are asymmetric faults, commonly associated with insulation failures, caused by various causes, such as poor connection, overloading and overheating. The types of stator faults that are allocated to the threephase (3-ph) windings are shown in Figure 1. Studies have shown that stator faults create waveforms of unbalanced flux that cause the engine to draw asymmetrical phase currents and unbalanced air gap flux. In addition, the fault-induced distortion can be due to the shifts in the air gap flux caused by the resistive and inductive shift through the stator windings. This field induces frequency components in the stator windings that modulate with the supply frequency, since the rotor magnetic field contains harmonics related to the rotor slots. Therefore, harmonics around the base frequency are produced and modulated by rotor slot harmonics when a fault occurs. Generally, short circuits occur between turns of one phase in stator windings, or between turns of two phases, or between turns of all phases. The types of stator faults that are allocated to the three-phase windings are shown in Figure 1. For the case of short circuits between turns in the same phase, we try to introduce only in the matrix the coefficients of resistance and inductance of the stator and in the mutual inductance stator-rotor. To model the short-circuit fault between the coils in an asynchronous machine, we follow this model (Taherzadeh et al., 2017) Us\u00bd = Rs\u00bd Is\u00bd + P[s\u00bd \u00f01\u00de 0\u00bd = \u00bdRR IR\u00bd + P[R\u00bd \u00f02\u00de [S\u00bd = Mss\u00bd + LSf IS\u00bd + MSR\u00bd + IR\u00bd \u00f03\u00de [R\u00bd = MRS\u00bd + MRR\u00bd + LRf IS R \u00f04\u00de With [Us]= usa usb usc 2 4 3 5, [Is]= isa isb isc 2 4 3 5, [ s]= usa usb usc 2 4 3 5 representing the voltages, the currents, and the stator flux respectively",
            " One distinguishes the direct effects, where the application of a magnetic field (or the variation of this field) gives a modification of a mechanical parameter. In the reverse effects, the variation of a parameter or a mechanical constant of a material gives a modification of its magnetization, where the initial embedding of this material in a magnetic field is not conditional. The two magnetic and mechanical domains are coupled thanks to this material (Mohammadi and Esfandiari, 2015). Modeling and simulation of a magnetostrictive cantilever Figure 1 shows a schematic diagram of energy recovery. The system is composed by a beam cantilever, a magnetic circuit and a magnetoelectric transducer. The latter is a sandwich of a PZT layer bonded between two layers of Terfenol-D. Magnetostrictive layers are longitudinally magnetized. The operating principle of the device is as follows (see Figure 2). A magnetic circuit displacement relative to the transducer is done when the harvester is excited (Figure 3). The ME transducer undergoes magnetic field variations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure2-1.png",
        "caption": "Fig. 2. Configuration of the proposed RPS-HTSFS machine: (a) Overall view and (b) Exploded view.",
        "texts": [
            " 1, in the conventional HTS-FS machine, the armature windings and HTS-excitation coils are placed on the same stator, namely, the HTS-excitation coil spans two adjacent 1051-8223 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 11:27:00 UTC from IEEE Xplore. Restrictions apply. stator teeth. The overlong end windings cause the waste of HTS wires, thus increasing the cost. This paper proposes a new 12-slot/10-pole HTS-FS machine as shown in Fig. 2. The key of the design is that the single ring-shaped HTS-excitation coil is independently placed in the coaxial claw-pole outer excitation stator while the armature windings are wound around the U-shaped iron core on the inner flux-switching armature stator without permanent magnets. That is, the radial partitioned stator structure consists of two parts, namely, inner armature stator and outer excitation stator. The inner armature stator is composed of 12 U-shaped irons wound by armature windings while the outer excitation stator has six claw-poles located on each side and every claw-pole radially connects the U-shaped inner stator iron yoke"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000508_robio49542.2019.8961531-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000508_robio49542.2019.8961531-Figure3-1.png",
        "caption": "Fig. 3. Samples with DART physics engine in the simulation, which show the grasp effect in both vertical and horizontal directions. We can grasp the same object in various ways.",
        "texts": [
            " Manual annotation and real-world acquisition can easily generalize to actual execution. However their drawbacks are higher cost, therefore not suitable for generating large-scale datasets. Methods based on analytical criteria or physical simulation can acquire a large amount of data in a short time, but analytical criterion only considers the capture moment, cannot measure the dynamic features after the execution. Therefore, we use the DART physics engine integrated in Gazebo simulation platform for data generation, it can easily verify the dynamic performance of the grasp. Figure 3 shows the results of several grasp attempts. In addition, DART simulation can also obtain the depth information of the object conveniently. Compared with the RGB image, depth information is more stable towards the changes of environment illumination or objects material, which seems appropriate for constructing the map from extracted feature to the grasp quality. Errors can come from everywhere, such as the noise of robotic arm\u2019s joints, the inaccuracy of camera and the robot\u2019s calibration. So uncertainty is added to improve generalization performance[23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001500_0954411920950904-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001500_0954411920950904-Figure2-1.png",
        "caption": "Figure 2. The designed 6-DOFs CT-guided surgical robotic system: (a) sketch diagram of clinical practice for percutaneous lung puncture, (b) description of brief operational flow, (c) 3D model of the designed robot, (d) enlarged view of orientation mechanism, and (e) real object. The numbers denote the following: 1-base, 2-supporting shield, 3-lift platform, 4-linear motion platform in left and right, 5-linear motion platform in front and back, 6-orientation mechanism, 7-force sensor, 8-clamp, 9-needle, 10-servomotors, 11-harmonic gear reducers, and 12-synchronous pulley.",
        "texts": [
            "6KPa; however, for larger strains of less than 90%, the yield stress syt is 116.9KPa, and when e \\ 0.8859, the compressive modulus Etc is a quadratic polynomial curve. Details for the s-e curve are presented in Appendix 1, Figure 10. The 6-DOFs CT-guided surgical robotic system is an experimental setup developed to perform the percutaneous lung puncture with accurate and consistent initial conditions, and to measure the relevant parameters (e.g. the puncture force) during insertion events, as shown in Figure 2(a). Figure 2(b) depicts the brief operational flow. First of all, the physician is able to get the information of the lesion recognition from CTguided images. Second, the control command is sent through the UI (User Interface) and the physical movement of the assisted robot is realized by the controller to complete the puncture positioning. Finally, the insertion operation from the initial insertion point to the lesion is guided and implemented by the physician with the aid of designed robot. During this period, the force sensor with a range of 200N and 10N m (HEX-E A2, a product of OnRobot Corp., Denmark) collects the puncture force data in real time. On one hand, it monitors the abnormal puncture force to ensure the safety of invasive surgery; and on the other hand, the feedback data can be used for the force-position control algorithm to increase the accuracy of robot-assisted insertion. The three-dimensional (3D) model of the surgical assisted robot is shown in Figure 2(c). It mainly consists of three parts which are the linear motion platform, the orientation mechanism, and the insertion needle. As the names imply, the linear motion platform, a rectangular coordinate type, is exhibited as a linear path along three directions of X, Y, and Z, and the orientation mechanism, an intercross type, as shown in Figure 2(d), is exhibited about a rotational path along three directions of a, b, and g. The former is used to adjust the position of the needle, and the latter is used to realize the rotational movement of the needle. Motion control of both can make the needle accurately reach the initial entry point and adjust the posture of the needle around the needle tip. The position and posture accuracy are within 0.4mm and 0.2 , respectively, and the repositioning accuracy is within 0.03mm, indicating that it has good control accuracy and can well satisfy the requirements for use. The insertion needle matched with a clamp is connected to the force sensor and mounted together to the end of the robot. All parts are attached to the base by connecting the supporting shield. Figure 2(e) shows the real object of designed surgical assisted robot. The relevant parameters of performance are expressed as follows: X-axis working range of 440mm, 400mm for Y-axis, 470mm for Z-axis, 360 continuous rotation for a-axis, b-axis of680 , g-axis of \u2013120 ;50 , 0.4m/s maximum working velocity of linear motion, 144 /s maximum angular velocity for rotational axes, and6 0.05mm positioning accuracy of designed robot. Moreover, the end-effector can also contain a redundant DOF in the direction of insertion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002629_012013-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002629_012013-Figure1-1.png",
        "caption": "Figure 1. 3D modelling of an electric motor",
        "texts": [
            " This arrangement permits the transfer of data from the planning stage into the stage of coming up with the manufacture of a product, without the requirement to reenter the data on part geometry manually. The info developed throughout CAD is stored; then it is processed any, by CAM into mandatory knowledge and directions for operative and dominant production machinery, material handling instrumentation, and automatic testing and review for product quality. The modeling of the electrical motor is done in 3D and it is shown in figure 1. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Methodology is that the systematic, theoretical analysis of the ways applied to a field of study. It includes the theoretical analysis of the body of ways and principles related to a branch of information. Typically, it encompasses ideas like paradigm, theoretical model, phases and quantitative or qualitative techniques"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001218_ec-10-2019-0447-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001218_ec-10-2019-0447-Figure8-1.png",
        "caption": "Figure 8. Processing map for peak temperature for laser power",
        "texts": [
            " The criterion for finalizing the NN architecture is based on the minimum mean square error (MSE). The optimal NN configuration for predicting peak temperature was found to consist of 48 hidden layer of neurons and 241 degrees of freedom of the network. The NN model architecture is shown in Figure 6. The trained network reaches the minimum MSE of 3.26 10 4. The Comparison between peak temperatures obtained from the FPM simulations and predicted by NN is given in Figure 7 showing the good prediction of the metamodel with a regression coefficient R2 = 0.9969. Figure 8 shows the typical processing map of peak temperature obtained using the built-up surrogate model for laser power Plaser = 400W and Plaser = 600W. All the process conditions obtained in the previous section are not suitable due to the criteria for the stable condition of the single clad track. The surrogate models built up in previous section were first used to predict 4,900 results of the peak temperature, clad angle and Figure 5. Three-dimensional and two-dimensional locations of powder feeding rate (Dm), laser power (Plaser) and scanning speed (Vs) combinations in the design space Figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000534_iros40897.2019.8968016-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000534_iros40897.2019.8968016-Figure12-1.png",
        "caption": "Fig. 12. Transmitter displaces with 90\u00b0 angle in respect with Receiver 1 and Receiver 2 axis.",
        "texts": [
            " Each experiment was performed with the use of the setup described in Section IV-A, the only difference being that in this set of experiments a second receiver was used. The oscilloscope\u2019s second channel was used in order to capture the Receiver 2 ZIF voltage signal. The transmitter displacement in reference to Receiver 1 is always 8 cm, while its displacement in reference to Receiver 2 is calculated subtracting transmitter\u2019s initial position from its final one. All distances between transmitter and Receiver 2 shown in Fig. 12\u201314 were calculated using the Pythagorean theorem for orthogonal triangles. In all three sets of results, the LOWESS curve-fitting method [36] was applied in order to be able to manipulate the waveforms and acquire accurate measurement results. Ten measurements were performed for each of the three sets of experimental data and the results for each one of those are shown in Table I. As elaborated in [29], it is once again assumed that a ZIF signal semiperiod (time between two consecutive zero crossings) correspond to 17 mm of transmitter displacement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001609_0021998320960531-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001609_0021998320960531-Figure1-1.png",
        "caption": "Figure 1. Illustration of SrPP structures manufacturing process. (a) Mould and sheet are heated up to required temperature and then required pressure was applied (b) Mould is precisely translated to the calculated step while keeping sheet at fixed position and process was repeated (c) Corrugated SrPP sheets that act as core (d) Profiles are passed through rollers to ensure dimensional quality. (e) Polymer adhesives were used to join FS with core (f) Unit cell and nomenclature.",
        "texts": [
            " So in order to overcome these problems we proposed ex-situ consolidation process, where the raw material SrPP was supplied in laminated form that assures the proper fibre/matrix contents to exhibit best structural properties. The basic principle was forming process at elevated temperature that is used to fabricate corrugated metal sheets. In order to deal with high shear and material flow problems that occur during conventional forming process; a steel mould was used with high precision step up translation capability in an elevated temperature chamber. The fabrication process for SrPP sandwich beams (SrPPSB) is schematically shown in Figure 1(a) to (d). Corrugated sheets manufactured by forming process act as core of structures and then above mentioned polymers adhesives were used to join with face sheets. It should be noted here that the forming process was done at a temperature that do not affect the fiber/ matrix contents as well as maintain the structural integrity. According to Prosser et al. the basic criteria was the temperature at which material is soft enough for bending but lower than its melting point 165 C so no shearing or inter-toe damage occurs.23 We fabricated structures in temperature range 145\u2013160 C and 155 C was selected as optimal bending temperature because structures exhibited maximum compressive strength at this temperature while maintaining the fibre/matrix contents. After the required 155 C temperature was maintained the mould pressure was applied and sheet is bend into required profile as shown in Figure 1(a). After shaping the first profile from the laminated sheet, the mould was step forward to the specified distance while the corrugated profile was constrained to move towards mould during the next operation; Figure 1(b). The corrugated profiles were passed through heated rollers so that dimensions of all profiles are uniform as shown in Figure 1(c). Finally the face sheets were joined adhesively to the upper and lower horizontal core faces with same family of polymer adhesives so that criteria of maximum recyclability can be achieved as shown in Figure 1(e). As PP is a non-polar polymer so it is little bit challenging to achieve an intact bond between face sheets and core faces. The core faces and corresponding areas on FS were abraded to generate a fibrous surface that can enhance the adhesion through the phenomenon of mechanical locking.24 The fabricated structures are shown in Figure 2(a) to (d). It was observed microscopically that at 155 C bending temperature there was good inter-toe adhesion at bent corners (inset image in Figure 2(a)); at the same time fibre/matrix ratio was also retained at optimal level as shown in Figure 2(d)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.5-1.png",
        "caption": "Fig. 34.5 Structural design of the tail",
        "texts": [
            " The formers are circular in shape and have a maximum diameter of 172 mm, and the length of the fuselage including the boom is 1210 mm. A boom passes through from the fourth former onwards and is permanently fixed to the fuselage. All formers are made of aeroply, and the stringers are made of balsa. 3 mm thick balsa sheets and aeroply sheets of varying thickness are laser cut into stringers and formers, assembled using a rail and slot mechanism and joined using cyanoacrylate. The load bearing formers that hold the landing gear, and motor-mount were reinforced by using a thicker aeroply sheet. Figure 34.5 shows the assembled empennage. The horizontal stabilizer and vertical stabilizer have a total of 8 and 5 ribs, respectively. Both horizontal and vertical stabilizer have a balsa I section passing through C/4 of each rib, a 10 mm leading edge and a 3 mm web placed at a position where elevators and rudder begin, to add hinges. The rudder has a 30\u00b0 cut to provide maximum deflection to the elevator. Further, a fixture provided below the horizontal stabilizer facilitates easy attachment of the boom and empennage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002170_s00170-020-06399-z-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002170_s00170-020-06399-z-Figure15-1.png",
        "caption": "Fig. 15 a\u2013c The effect on distortion during LSP forming after 1, 5, and 9 shots, respectively, at 300 K. d\u2013f The effect on distortion after 1, 5, and 9 shots, respectively, at 500 K. Note that the images are taken from simulations that apply the LSP parameters of Case 4 in Table 4. (Each plot has a unique vertical axis scale)",
        "texts": [
            " Given that the maximum relative difference is within 6% at all the temperatures, the far simpler and convenient Gaussian profile does not go discredited. Note also that, in general, all cases show a similar trend and reveal relatively little variation in the fraction of conforming surface area at the various temperatures. Nonetheless, the optimum temperature for LSP forming of the 316L part is clearly 300 K. To better understand the manner in which the surface is \u201ccorrected\u201d when the LSP forming is conducted at different temperatures, Fig. 15 shows the surface topography at intermediate \u201cshot frames\u201d during the LSP forming process at 300 K and 500 K. Figure 15a, b, and c illustrate, respectively, the surface topography after 1, 5, and 9 shots at 300 K. Similarly, Fig. 15d, e, and f show the surface topography after 1, 5, and 9 shots at 500 K. Comparing Fig. 15a and d, 78.56% and 77.49% of the surface area meets the 2-\u03bcm conformance criterion after 1 shot at 300 K and 500 K, respectively. Similarly, comparing Fig. 15b and e, 85.19% and 75.13% of the surface area conform after 5 shots at 300 K and 500 K, respectively. Finally, with all 9 shots executed, Fig. 15c and d show, respectively, that 84.75% and 74.25% of the surface area conform at 300 K and 500 K. It is also noteworthy that the surface areas containing the highest peaks (visible as dark red contours) reduce during each respective multi-shot LSP treatment, i.e., from Fig. 15a to c at 300 K, and from Fig. 15d to e at 500 K. While this effect is more significant when the forming environment is 500 K (see Fig. 15d to f), it is also observed that the height of the tallest peak increases by \u223c 2 \u03bcm despite the increase in the conforming surface area. When comparing roughness metrics at the different stages of LSP forming at 300 K and 500 K, note that both average roughness (Ra) and root-mean-square roughness (Rq) reduce, respectively, from 1.5789 to 1.8846 \u03bcm (postSLM, Fig. 11a) to the values indicated in Fig. 15. These metrics predict slightly greater surface roughness at the higher forming environment temperature. Note that the values of roughness range (Rt) do not provide a consistent metric from which to assess the fraction of conforming surface area. This is because Rt values simply indicate the difference between the highest and lowest data points, disregarding the actual domain that is made to conform as the LSP forming treatment is executed. As discussed earlier in the LSP forming strategy in Section 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000170_ecce.2019.8913254-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000170_ecce.2019.8913254-Figure3-1.png",
        "caption": "Fig. 3. Case 1- Study Domain",
        "texts": [
            " For instance, the stator core tooth is subjected to saturation, the shape of the rotor pole tip will cause non-uniform flux distribution, and damper bar currents are not uniform on all damper bars. The question becomes: how much do those factors affect the bar forces in the slot in comparison with the simplified and conservative analytical calculation? A 2D time stepped electromagnetic finite element analysis was performed to investigate the contribution of those factors on the forces. The software Flux 2D from Altair was used for this investigation. The study domain was modeled and consisted of a section with 7 poles (Fig. 3). The sides of the model are subjected to periodicity boundary condition with anticyclic conditions. A relatively fine mesh was used, in particular on the damper bars and slot wedges. The total number of nodes exceeded 295,000 using with 2nd order finite elements. Fig. 4 shows a detail of the mesh around a part of the airgap. External circuit connections were used on the stator winding, field winding and damper winding. Each winding was represented with its inductance and resistance. Several iterations were run to calibrate the magnetic properties of the stator and rotor material such that the voltage induced at the generator terminal at several field current and saturation level match the actual open circuit saturation curve of the generator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001558_s10778-020-01007-9-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001558_s10778-020-01007-9-Figure5-1.png",
        "caption": "Fig. 5",
        "texts": [
            " To assess the practical efficiency of the found optimal laws of a crane with a load on a flexible suspension, experiments were conducted in a laboratory. 3. Conducting Laboratory Experimental Studies and Analyzing the Obtained Data. The purpose of the laboratory experimental studies was to establish the main quality indicators for the implementation of optimal control of the crane motion model when using the frequency control of its asynchronous electric drive. We used a laboratory setup and equipment that are shown in Fig. 5. The design of the equipment allows us to study the optimal motion control of the crane model with a load on a flexible suspension when using a frequency-controlled drive. The main parameters of the laboratory equipment are listed in Table 3.1. In the experiment, we used a Mitsubishi Electric FR-E700 frequency converter that controlled the angular speed of the crane model electric motor. A RaspberryPi microcomputer was connected to the frequency converter. Its software was developed for conducting experimental research: it was used to calculate and send control signals to the COM port of the frequency converter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001861_1350650120972499-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001861_1350650120972499-Figure1-1.png",
        "caption": "Figure 1. General structure of an axial piston pump.",
        "texts": [
            " When the texture radius and depth are set to 18 mm and 0.8mm, there exists the greatest load carrying force and lowest friction coefficient. This work presents a key designing guide for axial piston pump textured slipper bearings. Surface texture, axial piston pump, slipper bearings, spherical dimples, genetic optimization Date received: 8 June 2020; accepted: 19 October 2020 Axial piston pumps are commonly employed in various hydraulic systems, including construction, agriculture, and aerospace market segments. Figure 1 shows an axial piston pump whose rotating kit is consisted of a cylinder block, a valve plate, pistons, piston bushing (guide), and slippers (shoes) on the end of the pistons angled on a swashplate. Particularly, the slipper bearing is considered as the supporting part. The frictional performance of the slipper bearing is closely related to service performance and life cycle of the whole machine. With the development of mechanical etching and laser technology, the surface of the slipper bearing can been machined with dimples of appropriate size"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000008_aim.2019.8868551-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000008_aim.2019.8868551-Figure4-1.png",
        "caption": "Fig. 4. RSR versus RSL maneuver",
        "texts": [
            " However, for every, there exists a unique optimal global maneuver along the sequence among the 2n possible maneuvers (also proven in [14]). This can be demonstrated by considering the Dubins maneuver candidates between a pair of regions, where the start and end do not belong to the initial and final circles, respectively, as shown in Fig. 3. At least two maneuvers, LSL and RSL, can be eliminated since they result in arc lengths greater than \u03c0, which hints that the unique maneuver can be solved for via elimination. Consider the RSR and RSL paths presented in Fig. 4, which correspond to the local maneuvers between two regions with the initial region in the clockwise orientation. Using the path parameterization presented in the previous section, the difference between LRSR and LRSL can be expressed as: (4L)RSR\u2212RSL = A\u2212 2rregion\u03b8f , (13) where A = Lc \u2212 Ls \u2212 2r cos\u22121(LS LC ) 2r . (14) Note that the expression in (13) does not depend on \u03b8i, since both maneuvers follow the same path between any arbitrary \u03b8i to \u03b8 = \u03b2 on the initial circle, where \u03b2 is the angle that the line connecting the initial region to the final region makes with the horizontal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002912_s12555-020-0049-x-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002912_s12555-020-0049-x-Figure3-1.png",
        "caption": "Fig. 3. ARP sit to stand position transformation (A) sitting position, (B, C) movement during sit to stand position transformation and (D) standing position.",
        "texts": [
            " The associated characteristic equation is as follows: [\u03bb + c][\u03bb 3 +(k1/\u03b5o)\u03bb 2 +(k2/\u03b5o)\u03bb +\u03b1 2 3 (k2/\u03b5o)]. (37) The coefficient relationships for the SPO are given by the following: k1 \u03b5o = 3\u03bbd , k2 \u03b5o = \u03bbd , \u03b13 = \u221a \u03bbd 3 , c = K \u03b5c = \u03bbd . (38) 4. SIMULATIONS AND RESULTS In this section, the simulations and simulation results of the designed controller implemented on the STS mechanism in both operating modes are presented. The shift of the system from a sitting to a standing position is shown in Fig. 3. A SolidWorks STS mechanism model was converted into a Simulink model, as shown in Fig. 4; this represents the actual hardware with actual dynamics. It is essential to validate the controller before it is implemented on the real hardware to preclude causing any system damage resulting from malfunction. The black block in the model represents the patient\u2019s mass. The position of the linear actuator was controlled. The maximum torque (T ) required for a linear-actuator motor to move the load vertically [35] is given by T = ( J+ mr2 e ) a r + r e mg (39) where J is the moment of inertia of the motor including the lead screw; m is the total allowable mass on the linear actuator; a, e, r, and g denote the acceleration of the load in either the up or down direction, the fractional efficiency of the lead screw, the transmission ratio (linear motion to angular motion) or translational displacement of the load with one turn of the screw, and gravitational force, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001320_012049-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001320_012049-Figure4-1.png",
        "caption": "Figure 4. The finite element model of (a) solid tire components and (b) contact boundary conditions.",
        "texts": [
            "5 ( 1)i i i n i ii 1 U= K J            (4) and ( 1/3) ,i iJ  1 2 3J    where i is the deviatoric principle stretches, J is the Jacobean determinant, K is the initial bulk modulus, and i , i are constants. The 3-D solid tire model was created using miscellaneous methods the 3D scanner and ComputerAided Design (CAD) in response to the geometry of KOMACHI series II solid tire, which has 6.00- 9.00 inch in size. Each component of the solid tire was divided by hexahedral elements. The rubber layer elements are presented by different colors, as shown in Figure 4(a). The internal, middle, tread layer and steel wires were assembled by 16,623, 8,448, 14,366 and 7,040 hexahedral elements, respectively. In Figure 4(b) there are finite element pouts that are connected by conjunct nodes regardless of the connection between the outer tread and the inner tread layer. Both of them were defined in the glue contact condition. The purpose of the experiment was to set up the finite element model of the solid tire to compress on a rigid flat plate according to the physical tire stiffness testing. The multi-point constraint (MPC) was defined on nodes at the inner surface to link an axis of solid tire model and to model the steel wheel, which was installed by the solid tire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001530_0954406220957369-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001530_0954406220957369-Figure9-1.png",
        "caption": "Figure 9. Pressure profile at bearing mid-plane z \u00bc 0:5; a \u00bc 0:7\u00f0 \u00de for a spin speed of (a) 4000 rpm, (b) 5000 rpm, (c) 6000 rpm, and (d) 7000 rpm.",
        "texts": [
            " It may be seen from Figures 6 and 7 that bearing coefficients depends strongly on the extent of pockets \u2018a\u2019. Additional data for bearing coefficients for different speeds are given in Table 2. The threshold speed of instability (xth) with different spin speed (xs) is obtained for different extent of pockets (a) and shown in Figure 8. A plot of pressure at mid plane of bearing ( z \u00bc 0:5) along the bearing surface for all shapes of pockets for four different speeds (4000 rpm, 5000 rpm, 6000 rpm, and 7000 rpm) is shown in Figure 9. Moreover, a complete three-dimensional pressure profile and corresponding pressure contours is shown in Figure 10. It may be seen from Figure 8 that with the provision of pocket in the bearing, the threshold speed of instability can be increased for all pocket shapes if the extent of pocket is greater than 60% of the total circumferential length. The reason is that the pressure generated in the bearing with pocket is spread into a larger area as can be seen in Figures 9 and 10. In Figure 9, it can be seen that pocketed bearings modifies the pressure profile significantly. In case of pockets pressure is distributed over a larger portion of the bearing as compared with bearing without pocket. Also, the peak is shifted towards the pocket boundary. This modifies the resultant bearing forces such that bearing forces will act to stabilize the journal motion. The bearing with rectangular pocket shows a maximum increase in stability limit speed (23%) when a \u00bc 0:7. It can also be seen in Figure 9 that with an increase in spin speed, the magnitude of pressure decreases, however the profile remains similar. A decrease in pressure is due to increase in minimum film thickness because the journal will tend to move near the bearing center with increasing speed. Figure 10 shows a 3D view and contour plot of the pressure profile for a speed of 7000 rpm. At other speeds, the nature of pressure profile will be qualitatively same as depicted in Figure 9. It is clear from Figure 10 that there is a change in pressure gradient along the pocket boundary. This is because a sudden change in film thickness will increase the pressure at the boundary. Up to this point, all the results were presented for a flexible rotor. However, the stability limit speed of the rotor can be further improved if it is rigid. In the case of rigid rotor, the equation of motion will be simplified as given in equation (28). Therefore, if the rotor is assumed to be rigid as for the case of short rotors where diameter to length ratio of the shaft is high, the stability limit speed should increase as compared to flexible rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure17-1.png",
        "caption": "Fig. 17. Displacement of curvature center in 4PBB.",
        "texts": [],
        "surrounding_texts": [
            "According to Fig. 1, five sets of parameters were designed to verify the model, see Table 1. The bearing type is 6004, which is a widely used bearing type in worm and gear application. The following is the theoretical calculation and comparative analysis of the load-displacement curve of each bearing design Table 1. Five types of 6004 bearing design parameters. Item Unit Design 1 (ordinary DGBB) Design 2 (ordinary 4PBB) Design 3 (small curvature DGBB) Design 4 (big clearance DGBB) Design 5 (big contact angle 4PBB) Ball dia. mm 6.35 6.35 6.35 6.35 6.35 Pitch dia. mm 31 31 31 31 31 Contact angle \u00b0 0 25 0 0 35 Inner curvature - 0.515 0.515 0.505 0.515 0.515 Outer curvature - 0.525 0.525 0.515 0.525 0.525 Ball qty. - 9 9 9 9 9 Radial clearance mm 0 0 0 0.01 0 Bearing type - DGBB 4PBB DGBB DGBB 4PBB Max. axial load N 1000 1000 1000 1000 1000 Max. radial load N 1000 1000 1000 1000 1000 Max. tilting load N.m 10 10 10 10 10 Arm length mm 30 30 30 30 30 in Table 1."
        ]
    },
    {
        "image_filename": "designv11_63_0000522_cencon47160.2019.8974832-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000522_cencon47160.2019.8974832-Figure1-1.png",
        "caption": "Fig. 1. Schematic layouts and photos of fully populated EOCB backplane variants: (a) SEPPLANE1 with single glass waveguide panel and connector layout and (b) stepped waveguide crossover section, (c) top view of embedded glass panel showing single waveguide illuminated with 635 nm light, (d) SEPPLANE2 with dual glass waveguide panels and connector layout.",
        "texts": [
            " Finally, we report the results of a comprehensive test and measurement regime, whereby PRBS 231\u20131 optical test data at both 1310 and 850 nm was conveyed along the backplane embedded optical waveguides through the pluggable connector system from various optical test sources and validated for both fiber-toboard and board-to-board optical connectivity at data rates up to 32 Gb/s per channel. In Section VI we discuss the technical challenges and future work to make these technologies commerically viable. In order to accommodate the interconnect topologies of target system enclosures in data center environments, the optical layout for the EOCBs contains waveguide groups with varied point-topoint geometries between edge or midboard optical connector interface points. As shown in Fig. 1, two EOCB backplane variants were designed: SEPPLANE1 contains one embedded glass waveguide panel [see Fig. 1(a)], while SEPPLANE2 contains two smaller glass waveguide panels [see Fig. 1(b)]. The purpose of the latter was to demonstrate the ability to laminate multiple glass panels into a single PCB layer, which would be a critical means in future of providing optical interconnect across larger, higher density backplanes, if individual glass waveguide panel sizes were constrained. The designs include exposed windows in the PCB material to show sections of the glass waveguide panel embedded within. The SEPPLANE1 and SEPPLANE2 layouts each consist of eight waveguide groups. Each group comprises a row of 12 waveguides with a center-to-center channel separation of 250 \u03bcm, which is compliant with conventional parallel optical fiber interfaces such as those based on MT standards"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001450_1464420720948541-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001450_1464420720948541-Figure2-1.png",
        "caption": "Figure 2. Single-field, P, (a) and multi-field, D, (b) disposition of specimens in the build platform according to recoating direction, RD.",
        "texts": [
            " The influence of scan strategy as well as hatch and contour spacing were analysed. In terms of scan strategy, both \u2018stripes\u2019 and \u2018chess\u2019 methodologies were tested with a fixed vector size of 15 mm and chess size of 5 mm, respectively. It is important to mention fixed parameters such as laser spot size (80 mm) and layer thickness (50 mm). The laser power was of 270 W for hatch and 200 W for contour operations. Scan speed was also distinct for hatch (800 mm/s) and contour operations (600 mm/s). All specimens had the same build orientation. As observed in Figure 2, cylindrical specimens of P series were disposed in the build plate diagonally to recoating direction, for uniform wear of the LPBF machine raking system. The same was not performed for the D specimens given the goal to study the interface and surrounding region between the two scan fields. Support lattice structures were used in order to achieve cylindrical-shaped specimens. In order to guarantee the suitability of the used recycled powder, its density, flowability and particle size distribution (PSD) were measured and compared with identical materials in literature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure9-1.png",
        "caption": "Fig. 9. Flux density distribution of conventional HTS-FS machine: (a) Excited only by HTS-excitation field and (b) Excited only by armature field.",
        "texts": [
            " The key design parameters of the proposed machine, compared with the conventional HTS-FS machine, are listed in Table I. The armature winding can produce the magnetic field when the armature current flows in, which can make a disturbance to the HTS-excitation critical field state. Moreover, the over strong armature magnetic field will increase the quench risk Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 11:27:00 UTC from IEEE Xplore. Restrictions apply. of HTS-excitation coil. Fig. 9(a) and Fig. 10(a) show the flux density distributions of the analyzed HTS-FS machines excited by only HTS-excitation field. It can be observed that their flux densities on the U-shaped core are both up to 2.0 T, which can take full use of the strong magnetic field characteristics of the HTS excitation. However, when the armature current flows in, the armature windings can produce the magnetic field, which can make a disturbance to the HTS-excitation critical field state. Moreover, the over strong armature magnetic field will increase the quench risk of HTS coils. Fig. 9(b) shows the flux density distribution of the conventional HTS-FS machine excited by only armature current. Since the armature windings and HTSexcitation coils are placed on the same stator, the magnetic fluxes are all closed by the U-shaped core, causing the HTS coils easily affected by the armature reaction. As a comparison, Fig. 10(b) shows the flux density distribution of the proposed RPS-HTSFS machine excited by only armature current. Thanks to partitioned stator structure, it can realize the physical isolation of armature windings and the HTS-excitation coil"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002396_0142331220987532-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002396_0142331220987532-Figure3-1.png",
        "caption": "Figure 3. Mechanical model and the equivalent electrical circuit (Dai et al., 2009).",
        "texts": [
            " Modeling and simulation of a magnetostrictive cantilever Figure 1 shows a schematic diagram of energy recovery. The system is composed by a beam cantilever, a magnetic circuit and a magnetoelectric transducer. The latter is a sandwich of a PZT layer bonded between two layers of Terfenol-D. Magnetostrictive layers are longitudinally magnetized. The operating principle of the device is as follows (see Figure 2). A magnetic circuit displacement relative to the transducer is done when the harvester is excited (Figure 3). The ME transducer undergoes magnetic field variations. Magnetostrictive layers are obtained by changing the magnetic field to generate stress, which is then transmitted to the PZT layer; electrical energy is generated immediately. The equation of the voltage generated in open circuit: V0 t\u00f0 \u00de= laV B t\u00f0 \u00de m0 \u00f025\u00de The equation of the load voltage V= laV H3RL 1 jwC +RL = laV B3RL WC m0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1+ wCRL\u00f0 \u00de2 q \u00f026\u00de The equation of the optimal resistance of the electric circuit Ropt = 1 wC = tC 1 n\u00f0 \u00de nSE 11 + 1 n\u00f0 \u00deSH 33 2pfwle33 n 1 K2 31 SE 11 + 1 n\u00f0 \u00deSH 33 \u00f027\u00de The electromechanical conversion of energy is located in the gap"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure15-1.png",
        "caption": "Figure 15. Fifth and sixth eigenmode (133.93 Hz and 144.12 Hz)",
        "texts": [],
        "surrounding_texts": [
            "Composite materials, finite element analysis, automotive engineering, vibrational behavior, design optimization Date received: 21 March 2021; accepted: 5 May 2021"
        ]
    },
    {
        "image_filename": "designv11_63_0002442_etcce51779.2020.9350871-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002442_etcce51779.2020.9350871-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the manipulator system in X-Y plane.",
        "texts": [
            " Section VI discusses the result and Section VII concludes the work presented in this paper. 20 20 E m er gi ng T ec hn ol og y in C om pu tin g, C om m un ic at io n an d El ec tr on ic s ( ET CC E) | /2 0/ $3 1. 00 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ ET CC E5 17 79 .2 02 0. 93 50 /20/$31.00 \u00a92020 IEEE 978-1-6654-1962-8 97 8- 1- 66 54 -1 96 2- 8 Authorized licensed use limited to: Carleton University. Downloaded on May 29,2021 at 04:11:00 UTC from IEEE Xplore. Restrictions apply. II. FLEXIBLE MANIPULATOR Schematic diagram of the flexible manipulator system is depicted in Fig. 1. The curve line in the diagram represents a single link of the manipulator. It has a flexible structure where the body of the link is bent if a force is applied. Applying a torque, \u03c4 on the rigid hub causes the flexible link to rotate about an angle of \u03b8 from axis- 0x location to a new x - axis location. Due to the flexible characteristic of the link, the body deforms its shape. This is also known as an elastic deflection, u of the flexible link at an arbitrary location on its body. The elastic characteristic of the link causes the manipulator to vibrate unstably and introduces a disturbance during motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003130_s11837-021-04789-6-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003130_s11837-021-04789-6-Figure1-1.png",
        "caption": "Fig. 1. (a) Schematic of jet mill; (b) related flow field distribution at a gas pressure of 0.7 MPa; and (c) diagram of gas velocity and pressure distribution from the nozzle to the chamber center as pointed out in (a).",
        "texts": [
            " 40 lm was sectioned, mounted on a micromanipulator by deposition of platinum (Pt), and further thinned for HRTEM characterization. In general, jet milling employs inert gas as a transport medium that is accelerated by a Laval nozzle to reach a supersonic speed, by which powders in the chamber are driven to achieve the grinding/modifying effect enabled by both interparticle and particle-to-chamber collisions. The jet milling process used the following steps in this study. First, the raw Ti powders were injected and then accelerated in the milling zone (Fig. 1a) by the high-speed N2 gas for grinding and modification. Subsequently, the ground powders were delivered to the transportation region by the updraft force. In this step, coarse particles underwent an itinerant trajectory governed by centrifugal force back to the milling zone for further grinding, while some ultrafine particles may be pulled to the upper outlet and classified out by the grading wheel (Fig. 1a) due to its minor gravity. In other words, the coarse powders should achieve more intensive modification/grinding compared with the fine powders that were probably classified into the secondary hopper with less milling. In this study, jet milling was performed up to 8 min to modify or grind the Ti powders with N2 gas (4N purity with an oxygen level below 3 ppm) at a pressure of 0.7 MPa and a rotating speed of 4000 rpm at room temperature. A computational fluid dynamics (CFD) based simulation could help unravel the basic physical mechanisms ruling the milling process.24,25 In this study, we present a set of CFD calculations on a simplified ideal model for milling geometry. Technical details for the solver and turbulence model used in this study can be found in the supplementary material. A 2D map of gas velocity magnitude either inside or near the gas nozzle is presented in Fig. 1b, and centralized velocity and pressure of the gas fluid associated are shown in Fig. 1c. As indicated in Fig. 1c, the gas fluid reaches sonic speed at the nozzle outlet, after which its maximum velocity becomes supersonic (> 500 m/s) a few millimeters away from the nozzle outlet as presented in Fig. 1a. At this critical point, however, the gas pressure decreases to the lowest value, close to 0.1 MPa. It is noted that the developing trend between gas velocity and pressure is contrary (Fig. 1c). This is because when the compressed gas in the nozzle enters into the chamber, the acoustic gas expands suddenly, concomitant with the significant increased velocity but at the expense of pressure. The continued gas expansion becomes weaker so that the variation of velocity as well as pressure appears moderate as shown in Fig. 1c. Figures 2 and 3 demonstrate the PSD and SEM images as well as the repose angles of the raw, classified fine, and jet-milled Ti powders, respectively. Table II shows the particle size, SSA, particle size span, and apparent density of powders before and after jet milling. Additionally, gas-atomized Ti6Al4V powder was selected for Ref. 26. It can be observed that the raw powders exhibit a wide particle size range, while the PSD becomes narrower for the classified fine and jet-milled powders, as reflected by the results of morphology, particle size, and its span in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001005_icarsc49921.2020.9096150-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001005_icarsc49921.2020.9096150-Figure1-1.png",
        "caption": "Fig. 1. Abstract Dynamics Models: (a) represents the Linear Inverted Pendulum Model (LIPM); (b) represents three-mass model.",
        "texts": [
            " Thus, it does not only need a powerful resource to implement, but also it is platformdependent. Linear Inverted Pendulum Model (LIPM) is one of the common dynamics models in the literature [2]. The popularity of this model is not only because of its linearity and simplicity but also its ability to generate a feasible, fast and efficient trajectory of the COM. This model describes the dynamics of a humanoid robot just by considering a single mass that is connected to the ground via a mass-less rod (see Fig. 1.(a)). In this model, the single mass is assumed to move along a horizontal plane and based on this assumption, the motion equations in sagittal and frontal planes are decoupled and independent. Several studies used this model to develop an online walking generator based on optimal control [3] and also linear MPC [4], [5], [6]. Several extensions of LIPM have been proposed to increase the accuracy of this model while keeping simplicity level [7], [8], [9]. The three-mass model is one of the extended versions of LIPM [10], [11], [7]. As it is shown in Fig. 1.(b), this model considers the masses of legs and the body to increase the accuracy of the modeling. It should be mentioned that to keep the model linear, the vertical motions of the masses are considered to be smooth and the vertical accelerations are neglected. In this paper, the three-mass dynamics model is used as our 978-1-7281-7078-7/20/$31.00 \u00a92020 IEEE 257 Authorized licensed use limited to: Murdoch University. Downloaded on June 16,2020 at 20:25:56 UTC from IEEE Xplore. Restrictions apply",
            " To do that, we assume that the COM of the robot is located at the middle of hip and, based on this assumption, the overall dynamics of the robot is firstly restricted into COM and then the reference trajectory for the hip will be generated using the analytical solution of the LIPM as follows [17]: ph(t) = pst + (pst\u2212phf ) sinh ( (t\u2212t0)\u03c9 ) +(ph0 \u2212pst) sinh ( (t\u2212tf )\u03c9 ) sinh((t0\u2212tf )\u03c9) , (12) where t0 and tf represent the beginning and the ending times of a step, ph0 , phf are the corresponding positions of COM at these times, respectively. After generating the ZMP and the hip trajectory, the swing trajectory should be planned. To have a smooth trajectory during lifting and landing of the swing leg, a Bezier curve is used to generate this trajectory according to the generated footsteps and a predefined swing height. The masses trajectories can be easily generated based on geometric relations between the generated hip, ZMP and swing leg trajectories. According to Fig.1(b), the mass of the stance leg is located in the middle of the line between the ZMP and the hip. Similarly, the mass of the swing leg is located in the middle of the line between the swing foot and the hip. To show the performance of the planners presented in this section, an exemplary path planning scenario has been set up, which is depicted in Fig. 3. In this scenario, the robot stands at the START point and wants to reach to the GOAL point. The planning process is started by generating an optimum obstacle-free path (black-dash line)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure9-1.png",
        "caption": "Fig. 9. Curvature center under pure radial displacement.",
        "texts": [
            " (9) and (10), the displacement xi\u03b4 of the i-th steel ball is ' 0 0ni A A\u03b4 = \u2212 ( )22 1 2 0 xid d A\u03b4= + + \u2212 . (12) The equation of normal contact deformation ni\u03b4 of the i-th steel ball under which the pure ring axial displacement x\u03b4 occurs is established. When bearing\u2019s outer ring has a pure radial displacement in y direction (see Fig. 8), the raceway at the azimuth position i\u03c6 will have a radial displacement yi\u03b4 which can be expressed as, cosyi y i\u03b4 \u03b4 \u03c6= . (13) When the raceway has a radial displacement yi\u03b4 , the rela- tionship of the raceway curvature center is shown in Fig. 9. The curvature center distance ' 0A after deformation is, ' 2 ' 2 0 1 2 A d d= + ( )2 2 1 2 yid d\u03b4= + + . (14) The normal contact deformation between the steel ball and the raceway is, ' 0 0 ni A A\u03b4 = \u2212 ( )2 2 1 2 0yid d A\u03b4= + + \u2212 (15) ( )2 2 1 2 0y id cos d A\u03b4 \u03c6= + + \u2212 . Here an expression for the normal contact deformation ni\u03b4 under pure ring radial displacement y\u03b4 is established. Similarly, when the outer ring displaced z\u03b4 in z direction, the normal contact deformation ni\u03b4 is expressed as ' 0 0ni A A\u03b4 = \u2212 ( )2 2 1 2 0 sinz id d A\u03b4 \u03c6= \u2212 + \u2212 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002604_j.apacoust.2021.108063-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002604_j.apacoust.2021.108063-Figure13-1.png",
        "caption": "Fig. 13. Modal experiment: (a) the arrangement of sensors and (b) modal results of the rear door.",
        "texts": [
            " The effectiveness of the equivalent model was validated by the modal and vibration transfer function (VTF) experiments. Moreover, utilising the equivalent model, the influence of the pre-compression on VTF from the chassis connecting points to the target point on the door was investigated. These results can be used as guidelines for dealing with the door vibration problem. To demonstrate the effectiveness of the proposed model, a modal experiment was firstly conducted to obtain the natural frequency and mode shape of the door with the door closing. The modal experiment is shown in Fig. 13(a). It was finished on LMS Test. Lab. A hammer with an appropriate tip was used as an exciter to invoke the resonance frequencies. For the decrease of leakage and noise, a force-exponential window was used in the input force channel and an exponential window was adopted in the response channels. An H1 estimator was utilized for calculated the frequency response function. In addition, to obtain a reliable result, every point was excited by five times. The force signal and the acceleration signal at measured points were recorded. The least-squares complex frequency-domain method was used for modal parameter identification. The identified first-order mode frequency of the door is 26.4 Hz, and the corresponding mode shape is depicted in Fig. 13(b). In the numerical analysis process, a trim-body instead of a full-scale vehicle was employed to analyse the door mode, considering that the chassis structure has nearly no influence on the door mode. Besides, adopting a trim-body allows reducing the computational time. In the simulation process, the spring model and the equivalent model were employed respectively. In the spring model, the stiffness of the seal strip was held constant at 5 N/mm, which is an empirical value. For the equivalent model, the dynamic stiffness was obtained from an FE analysis at a 4 mm pre-compression, which was experimentally determined by conducting a seal strip gap test when the door was locked. The integration of the equivalent model into the trim-body model was implemented by CBUSH and CVISC elements in the Hypermesh software. The comparison of the experimental and simulated results is shown in Fig. 13(b) and Fig. 14. It is observed that the mode shapes derived from experiment and simulation are similar, and the errors of the mode frequency with the experimental value are 0.7 Hz in the spring model and 0.9 Hz in the proposed model. From the modal analysis standpoint, both the spring model and the equivalent model can be utilised to represent the seal strip. VTF is a crucial index in the exploration of vehicle body vibration characteristics, which allows the designer finding the most sensitive excitation points in the vibration transmission",
            " (6): Gff f\u00f0 \u00de \u00f06\u00de where Gfa(f) is cross-spectral function of input and output; Gff(f) represents self-spectrum function of input. Thus, the VTF can be easily obtained using experimental method. For further validation of the proposed equivalent model, the experimental and simulated VTF results are compared, as illustrated in Fig. 15. The simulated VTF was calculated using FRF analysis in Nastran Solver and the experimental VTF was obtained using impulse hammer excitations based on the LMS Test. Lab. All the VTFs from chassis connecting points to target point S3 on the door (S3, as depicted in Fig. 13, on which the largest vibration acceleration was found) were measured and numerically calculated. However, Fig. 15 only shows the comparison of X-direction VTF from the chassis rear stabilizer bar connecting point to target point, considering that a similar result can be concluded at VTFs of other points and directions. In addition, the selected VTF is the largest in all the measured VTFs in the first resonance frequency region (in the vicinity of 26.5 Hz), which draws much attention to this studied door vibration problem",
            " The door vibration problem investigated in this study is described as follows: in the vehicle development of the prototype test stage, the vehicle presented an unacceptable vibration on the rear door when driving on a washboard pavement at the speed range from 11 km/h to 15 km/h. However, it performed well at other speeds or on other types of standard test pavements. Therefore, it can conclude that the source excitation of the problem is not from the engine. The excitation frequency was identified from 20.1 Hz to 27.4 Hz using the washboard pavement size and the speed range. In addition, a door modal experiment was conducted, and the identified mode frequency was 26.4 Hz, as shown in Fig. 13. Thus, this problem can be classified as a modal resonance problem. However, in this development stage, the modification of the door structure would incur a high cost for changing moulds. Therefore, alternative strategies are needed. In the current study, the effect of pre-compression on the VTF was studied. Based on the analysed results, the amount of pre-compression of the door seal strip was redesigned. The equivalent model is capable of analysing the vibration transmission through the seal strip at different levels of the precompression"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003416_tmag.2020.3020126-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003416_tmag.2020.3020126-Figure1-1.png",
        "caption": "Fig. 1. Configuration of the proposed PHE-HSM [8].",
        "texts": [
            " Although the flux regulation principle has been briefly explained by the harmonic analysis in [8], the mathematic analysis for the air-gap harmonic effectiveness is not given while the dominant harmonics and torque contributions of harmonics are not investigated. In this paper, the field modulation effect of the proposed PHE-HSM is analyzed, which can provide insight into the flux regulation principle, harmonic effectiveness, dominant harmonics, and torque contributions of harmonics. The field modulation effect of the proposed PHE-HSM is theoretically analyzed and verified by finite-element analysis (FEA) and experimental results. II. MACHINE CONFIGURATION Fig. 1 shows the proposed PHE-HSM including the machine topology and armature winding connection. The proposed machine consists of a 24-slot inner stator that includes the steel lamination, armature windings, PMs and DC-field windings, and a 14-pole outer rotor that is simply the steel lamination. the field windings are connected with a simple DC winding connection as shown in Fig. 1(a) while the armature windings are accommodated in the slots with concentrated windings as shown in Fig. 1(b). The distinctive feature of the proposed machine is that the DC-field pole-pairs are shifted from the PM pole-pairs to achieve the parallel-hybrid excitation. The H Authorized licensed use limited to: Carleton University. Downloaded on September 19,2020 at 20:32:59 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"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001691_icuas48674.2020.9213833-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001691_icuas48674.2020.9213833-Figure1-1.png",
        "caption": "Fig. 1: Parameters for the proposed approach: Efficient packing for full coverage of the area by deployment of synchronised homogeneous UAVs. The loiter circles and their instantaneous phase and coverage footprint is shown for the two presented cases: (a) Square packing; (b) Hexagon packing.",
        "texts": [
            " al [19] studied a facility location problem for groups of Dubin\u2019s vehicles, non-holonomic vehicles constrained to move along planar paths of bounded curvature, without reversing direction. Given a compact region and a group of Dubin\u2019s vehicles, the coverage problem is to minimize the worst-case traveling distance. This work presents an algorithm for persistent coverage of an area, by patrolling with a fleet of loitering fixed-wing UAVs at a pre-specified altitude over the area. The loitering circles are packed by inscribing over the packed squares or hexagons as shown in Fig. 1. In this paper, the circles are packed for both the cases over a rectangular area only. The user is supposed to provide with the number of available UAVs and the desired loitering altitude. It is desired to have enough number of UAVs to be able to start at minimum loitering circle, and have persistent coverage of the area at all times. Based on the available UAV count, the algorithm formulates and solves an optimization problem to compute the centre location of the uniform loitering circles for the given altitude, by maximizing the loitering radius",
            " Finally, Section V lists some of the possible future work in this domain and concludes the paper. The area to be covered by a UAV fleet can be represented as a graph with the vertices representing the locations of the agents (longitude and latitude). The deployed UAVs would represent a different set of nodes V\u2032 as a sub-graph G\u2032 = (V\u2032,E\u2032), and the virtual edges (E\u2032) between the neighboring UAVs represent the active communication link. We define the following quantities to facilitate the presentation of the proposed method. Fig. 1 summarizes these quantities graphically. Field of view (FOV): The FOV is a physical property of the sensor being used by the UAV and defines the coverage footprint based on the altitude of the platform. In Fig. 1, The FOV has been marked by \u03b8. Based on the sensor used, the sensing quality (q) can be defined as q \u221d 1/h, where h is the loitering altitude of the UAV. This means that the coverage quality decreases linearly as the altitude increases, and vice-versa. Coverage radius (rc): The coverage radius is the radius of the coverage footprint of the on-board sensor, given the sensor FOV and the instantaneous height (h) of the vehicle. 1232 Authorized licensed use limited to: Middlesex University. Downloaded on October 18,2020 at 17:34:33 UTC from IEEE Xplore. Restrictions apply. Coverage radius is directly proportional to the loitering altitude and inversely proportional to the coverage quality, for a given FOV. As seen from Fig. 1, it is defined as rc = h \u00b7 tan \u03b8. Loiter radius (rl): Any form of a fixed-wing vehicle has constraints on maneuverability and it cannot stay stationary while it is airborne. Instead, it can fly in a circle over the region of interest, called the loiter circle. The radius of that circle at a given instant is called the loiter radius. In Fig. 1, the loiter radius has been shown by rl. The physical properties and cruising velocity of a fixedwing UAV system define the lowest value of the loiter radius, called the minimum turning radius [20], given by, rmin-turn = v2 g \u03c8max, where v \u2208 R2 is the horizontal vehicle velocity, \u03c8max denotes the maximum bank angle and g denotes the gravitational acceleration. It is desired to have rl > rmin-turn to be able to provide coverage while causing less physical strain on the UAV. Maximum loiter (rl-max): This is the maximum loiter radius at which the UAVs can fly, while maintaining full coverage of the area",
            " Full coverage: It is defined as the state when each point in the area is guaranteed to be covered by at least one of the loitering UAVs at least once every loiter cycle during the operation. For N UAVs deployed in the area A, it is achieved when A \u2286 N\u2211 i=1 Ei. This serves as the main objective of the presented work, where we adjust the radius of the loiter circle for the available UAVs to achieve full coverage. Phase synchronization: For a UAV loitering at an altitude, the phase has been defined in this paper as the angle at which they are. It has been shown in Fig. 1 as \u03c6. We assume that all the loitering UAVs have the same phase at every instant of time for maximum separation, and hence the largest effective coverage. Super-agent: This is an agent with enhanced communication and computation capability, which is used as a global planner in case of simultaneous multiple node failures. 1233 Authorized licensed use limited to: Middlesex University. Downloaded on October 18,2020 at 17:34:33 UTC from IEEE Xplore. Restrictions apply. -100 -50 0 50 100 150 200 250 -100 -50 0 50 100 150 200 250 1 2 3 4 5 6 7 Fig",
            " Also, the UAVs are often equipped with efficient inertial measurement unit (IMU) and global positioning system (GPS) sensors for accurate location, altitude and orientation information (accurate to few centimeters). For real life situations, the assumptions like same cruising velocity, always synchronized phase, lagfree communication may pose obstacles like collisions and data package drops. Relaxing these assumptions will serve to make the algorithm more suited to practical applications, versatile, and scalable, which is among the future scope of this work. The details of the proposed approach are presented in this section. Fig. 1 shows the basic set up for the approach. In Fig. 1(a), the UAV is shown loitering at an altitude h over a square packed area, along the loiter circle with radius rl with instantaneous coverage footprint marked by rc. The sensor FOV for the given altitude has been marked by \u03b8. Fig. 1(b) shows the equivalent setup and parameters for the hexagonal packing. The proposed algorithm deploys the UAV fleet over the area with either of those packing methods and handles the cases of simultaneous multiple UAV failures, as summarized in Algorithm 1. The upper bound on the run time of this algorithm is O(N), for a network of N UAVs. The details of the approach have been discussed below. A. Initial Deployment This phase of the algorithm deals with the initial deployment of the UAVs in the area, based on the available UAV count and loitering altitude, by using location optimization technique to calculate the loiter radius and the locations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000145_jfm.2019.780-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000145_jfm.2019.780-Figure2-1.png",
        "caption": "FIGURE 2. Sketch of a spherical stick\u2013slip Janus particle and the coordinate system.",
        "texts": [
            " For an SSJP with a tiny stick portion, in particular, the force response can even display a re-entry Basset decay at high frequencies and then level off to a plateau at lower frequencies, showing a distinctive re-entrant history force transition (RHFT) that exists only for an SSJP. Below we demonstrate how we arrive at the above features by solving the oscillatory Stokes flow equation in \u00a7 2 and by using the matched asymptotic boundary layer theory in \u00a7 3. Consider the motion of a spherical SSJP of radius a in an incompressible viscous fluid of density \u03c1 and viscosity \u00b5. As depicted in figure 2, the slip part of the particle has polar angle \u03b80 in division with the remaining stick part. The particle undergoes an oscillatory translation at speed Upe\u2212i\u03c9t with the peak velocity Up and frequency \u03c9. It is more convenient to solve the problem in the translating coordinate system with the origin at the instantaneous centre of the particle. Having length, time, velocity and pressure scaled by a, \u03c9\u22121, Up and \u00b5Up/a, respectively, the velocity field v\u2032 and pressure p\u2032 around the particle are governed by the continuity equation and the unsteady Stokes flow equation, \u2207 \u00b7 v\u2032 = 0, (2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001102_j.matpr.2020.05.310-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001102_j.matpr.2020.05.310-Figure4-1.png",
        "caption": "Fig. 4. Deflection of the wheel housing.",
        "texts": [
            " The wheel housing designed was meshed in the software for proceeding with the nodal analysis to be performed in the ANSYS software [23]. The meshed view is shown in the Fig. 2. The mesh is then opened in the software for the analysis in which initially the load is applied to check for the stresses. The load applied gets enforced in the housing which is portrayed in the Fig. 3. On applying the stresses in the housing the deflection is noted in the output as 0.42 mm for the applied load as shown in the Fig. 4. The stress acting on the housing of the wheel while applying load is 39.008 N/mm2 as al., FEA based approach on replacing the metal cast wheel into thermoset 20.05.310 shown in the Fig. 5. The values determined were found to be within the safety limits since the maximum permissible value was 118 N/ mm2 mentioning consideration in usage of the made wheel housing. As an outset, considering various calculations and design development for the product and mould design for SMC, wheel housing was completed with actual values as guidance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002388_masy.202000247-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002388_masy.202000247-Figure1-1.png",
        "caption": "Figure 1. Shear specimen geometry: (a) Technical drawing including dimensions, (b) Principle for width correction for shear area by surface roughness Pz not to scale.",
        "texts": [
            "85 mm in diameter. The fiber weight content is approx. 12.5 wt.-% and the fiber diameter is 7 \u00b5m by a typically length of fibers between 150 and 400 \u00b5m after fabrication to filament. For the mechanical investigations according to DIN 65148 the specimen geometry has to be adapted. The basic principal, the characterization of the shear strength by a tensile load, is unchanged. The use of the original specimen geometry leads to an irregular failure mode. Therefore, the geometry is adjusted as shown in Figure 1a. The idealized shear area of the specimen is 6 mm \u00d7 6 mm. Due to the macroscopic waviness in the untreated \u201cas-built\u201d surface, the effective shear area is affected. Thus, themacroscopic dimension of the widthwwas determined with a tactile micrometer. This macroscopic result was corrected by the microscopically measured primary surface profile Pz. For the length l the idealized value of 6 mm is assumed. Considering this, the shear area is calculated following Equation 1.[15] Aeffective = (wmacroscopic \u2212 (2 \u00d7 Pz)) \u00d7 l (1) The results of the non-destructive testing methods for quality assessment are plotted in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure5-1.png",
        "caption": "Fig. 5. Machine flux line distribution of the proposed HEBFM at (a) IDC>0, (b) IDC=0, and (c) IDC<0.",
        "texts": [
            " \u03a6fA, \u03a6mA and \u03a6sA represent the amplitude of flux linkage generated by field, by PM and total flux linkage in Phase A respectively. It can be seen from Fig. 4 (a) and (c) that the amplitudes of the flux linkages produced by PMs and DC currents are invariant but the angles between the two vectors are changed. Consequently, the winding factor modulation coefficient is changed, which leads to the variations of total Back-EMF. In a word, the Back-EMF regulation of HEBFM is not implemented by magnetic field weakening or strengthening but winding factor modulation. Fig. 5 shows the machine flux line distribution in different status. It can be seen that with field current injected, the flux lines pass through every neighbouring slots. When there is no field current employed, the flux lines only pass through the slots with PM segments. The results clearly verify the analysis of flux path. Authorized licensed use limited to: Carleton University. Downloaded on June 05,2021 at 19:52:57 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003041_s10845-021-01803-1-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003041_s10845-021-01803-1-Figure1-1.png",
        "caption": "Fig. 1 One optical inspection system and the schematic diagram of its measurement",
        "texts": [
            " This inspection method significantly reduces the flow time in multistage manufacturing and improves the performance of the quality inspection. Before developing our task allocation and coordinated motion planning methodologies for multiple robots, we firstly introduce the physical modules of this autonomous multi-robot optical inspection system and its characteristics. These characteristics will have a large impact on methodology development. This section introduces how the optical inspection system works and what constraints it has. Figure\u00a01 shows one optical inspection system and the schematic diagram of its measurement. The non-contact probe uses an optical sensor based on the triangulation principle. The sensor projects a laser and a planar light simultaneously, which are used to obtain the accurate position of a measurement point (MP) in space according to the reflected light in the charge-coupled device camera imaging. Unlike 3D scanners that collect point cloud data offline, the new optical inspection gauge can measure the deviation of a single MP from each viewpoint of the probe in an inline setup"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002074_5.0031388-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002074_5.0031388-Figure7-1.png",
        "caption": "FIG. 7. Experimental setup for probing (a) zonal and (b) non-zonal modes. In either case, a balloon is vertically oscillated by a mechanical vibrator. (a) A pair of LED lamps illuminate the balloon\u2019s meridian, and a camera records the balloon\u2019s shape with long exposure time to capture the polar standing wave. A piece of sandpaper (SP) is used as the background to enhance contrast. (b) An LED lamp illuminates the oscillating balloon through a mirror on the top of the balloon. A camera captures the balloon\u2019s top-view snapshots through the same mirror.",
        "texts": [
            "0031388 32, 124113-6 Published under license by AIP Publishing no more water is injected into the balloon beyond the dot. The curve segments to the right of the dots are recorded to identify the pressure in a relaxed balloon. It is this pressure, not the pressure at the circular dot, that is used in subsequent analyses. Recall that this study examines the resonance of balloons that have not passed the second turning points of their P\u2013V curves. According to Fig. 6, this is the case for all balloons here. In experiments, balloons are mechanically oscillated along their axes of symmetry. As illustrated in Fig. 7, an inflated balloon is driven by a mechanical vibrator (model VTS-100 by Vibration Test Systems, USA) in the vertical (z) direction. The oscillation signal is a single-frequency sine wave. It is prescribed by a function generator (model AFG-3022 by Good Will Instrument Co. Ltd., Taiwan) and intensified by a power amplifier (model XTI-2002 by Crown Audio/HARMAN Professional, USA) before it is delivered to the vibrator. For each balloon, zonal and non-zonal modes are probed in two separate manual frequency scans. As illustrated in Fig. 7, each scan adopts a slightly different imaging method. For zonal modes, side-view imaging clearly reveals their shapes. In experiments, these modes are observed from the side with exposure times much longer than their periods of oscillation. As shown in Fig. 7(a), this method accumulates a balloon\u2019s polar waveform in the resulting image and enables clear identification of the standing wave pattern. The images of zonal modes are captured with a CCD camera (model 1200s by PCO AG., Germany) through a macro lens (model AF Micro-Nikkor 60 mm f/2.8D, by Nikon Corporation, Japan) at 10 frames per second (fps). The front view in Fig. 7(a) depicts part of the experimental setup viewed from the camera toward the balloon, i.e., in the x-direction. LED lamps are arranged to project light along the balloon\u2019s meridian for optimal visualization. A piece of sandpaper (SP) is used as the background to enhance contrast. As shown in the lower left in Fig. 7(a), the setup effectively illuminates the balloon\u2019s side lobes and exposes the features of the mode. For each balloon, the first five zonal modes are probed. The lowest possible forcing acceleration is applied. Once a zonal mode appears, the forcing frequency f 0 that triggers the maximum amplitude of the mode is pursued. Then, an image of the mode shape is recorded as already described. To estimate the frequency of the mode, 200 snapshots are captured with a frame rate greater than 2f 0. The frequency is then estimated based on the frame rate, the number of full oscillation periods, and the corresponding number of snapshots.79 In this study, all zonal modes are observed to oscillate harmonically, i.e., at the forcing frequency. Therefore, the frequency that triggers the maximum amplitude of a zonal mode is identified as the mode\u2019s resonance frequency. Probing non-zonal modes requires taking top-view snapshots of an oscillating balloon. As shown in Fig. 7(b), the balloon is illuminated with an LED lamp through a mirror. Sequences of top-view snapshots are then captured through the mirror with the same camera for zonal modes. As presented in Fig. 7(b), the simple setup successfully captures the triangular mode shape of an oscillating balloon. The resonance frequency of a non-zonal mode is estimated from its top-view snapshots with the same method for zonal modes. As reported later, several non-zonal modes are triggered with such high forcing frequencies that the camera cannot capture their snapshots at twice of these frequencies with a reasonable resolution for mode identification. In that case, the frame rate is reduced, and the resonance frequency of the mode cannot be estimated without aliasing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000679_edpc48408.2019.9012063-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000679_edpc48408.2019.9012063-Figure10-1.png",
        "caption": "Fig. 10 Wires in slot area after winding of pressed Sequences",
        "texts": [
            " Here a cut section through one side of several final manufactured Aluminum winding is shown. After the bending process, wires of different thicknesses lie on top of each other and lead to a trapezoidal shaped coil. The size of the cross-sectional areas of the single wires are almost the same. Currently, the focus of research is on the post-processing of the winding head. Due to the winding of the press sequences, the non-pressed round areas lie in the winding head of the coil, which means that the slot areas of the coil have no contact with each other, see Fig. 10. For a functional coil, it has to be ensured that the wire layers of the slot area contact each other. Several different possibilities will be explored. One is the option of subsequent pressing of the bending area in the winding head with a special tool. Furthermore, the lengths and bending geometries of the individual windings in the coil head area can vary in order to compensate for the difference in height between pressed areas and the initial round wire. V. SUMMARY AND OUTLOOK In this paper, some approaches to increase the filling factor were briefly described"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000934_s42835-020-00429-2-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000934_s42835-020-00429-2-Figure1-1.png",
        "caption": "Fig. 1 The concept of an equivalent circuit coupled with FEA",
        "texts": [
            " This method is disadvantageous for calculating the magnetizing reactance, which greatly influences the characteristics of the induction motor. The finite element method (FEM) is able to obtain the distributed characteristics of the motor, but it requires much computation time and memory. In addition, it is difficult to predict the power factor and efficiency using the method because the losses produced in the motor are not taken into account. As a result, in this study, both methods are combined to solve the problems of each method. Figure\u00a01 displays the concept of the equivalent circuit coupled with finite element analysis (FEA). A new challenge is directly calculating the * Sung-Il Kim kimsi@hoseo.edu 1 Rotating M/C Innovation Technology Team, Hyosung Corporation, Changwon\u00a051559, Korea 2 Electric Motor Research Center, Korea Electrotechnology Research Institute, Changwon\u00a051543, Korea 3 Department of\u00a0Electrical Engineering, Hoseo Univerity, Asan\u00a031499, Korea 1 3 main parameters of the circuit by FEM and estimating the characteristics from the circuit model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000041_012009-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000041_012009-Figure4-1.png",
        "caption": "Figure 4: Descriptive figure of the experiment setup",
        "texts": [],
        "surrounding_texts": [
            "A typical UAV mainly consists of four subsystems: 1) Autopilot, 2) Communication, 3) Sensors payload, 4) Actuation. The main paper aim is identifying corona discharge current interference effect on each UAV subsystem, and also identifying the risk of introducing a shielding solution with respect to changing air breakdown characteristics between valve hall towers. Hence, two experimental setups were prepared to evaluate the risks of corona discharge current interference and air breakdown voltage changing."
        ]
    },
    {
        "image_filename": "designv11_63_0000647_s11370-020-00316-9-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000647_s11370-020-00316-9-Figure10-1.png",
        "caption": "Fig. 10 Snapshots of the spot welding path planned by TRM algorithm",
        "texts": [
            " First, ten collision-free pathswere generated by an asymptotically optimal motion planner RRT*, which was run for 10 h. In the learning phase, the GMM-based sampling strategy was used to construct the teaching roadmaps. However, in the query phase, the posture of the body parts wasmodified randomly nearby. In this experiment, our method was compared with a variety of sampling-based algorithms fromOMPL, such as PRM, RRT, RRT-Connect, SBL, KPIECE and RRT*. The given planning time was 1 h, all experiments were performed ten times, and the average performance was recorded. Figure 10 is the snapshots of the spot welding path planned by the proposed TRM algorithm. As shown in Table 3, the proposed TRM planner outperforms the other sampling-based planners in terms of both success rate and planning time in this scenario. The experimental results show that the precomputed teaching roadmap can help to avoid unnecessary exploration and accelerate the motion planning process in a semistructured environment. This paper has introduced a teaching roadmap (TRM) algorithm for motion planning in a semistructured environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001439_j.optlastec.2020.106536-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001439_j.optlastec.2020.106536-Figure2-1.png",
        "caption": "Fig. 2. Photo of the experiment setup for line-spot-laser-actuated trapping and operation at liquid interfaces.",
        "texts": [
            " Then the surface tension difference can be calculated as the surface tension coefficient integral along the triple phase contact line. To better understand the line-spot-laser-based trapping and operation method, the numerical analysis in this work mainly focuses on three questions: how much surface tension difference can a line-spot laser generate; how much is the spacing between the line laser spots required for trapping; how does the halo ring size affect the operation force. The experiment setup for the line-spot-laser-actuated trapping and propulsion experiments at liquid interfaces are built as shown in Fig. 2, which mainly consists of a triple-axis electric slide table, a group of CW lasers and a container. The slide table is actuated by three step motors which are programmable and enable the slide table move automatically. The lasers are fixed to the slide table via a group of holders. Each laser holder has 4 degrees of freedom, through which the posture of the laser can be easily adjusted manually. Each laser is connected to a voltage adjusting module so that the laser power can be manually regulated, and the laser power is measured with the help of a power meter (VEGA, OPHIR)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002943_j.matcom.2021.05.034-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002943_j.matcom.2021.05.034-Figure1-1.png",
        "caption": "Fig. 1. Vertical rolling disc.",
        "texts": [
            "i=233562abb6. . Controllability in non-holonomic mechanics .1. Vertical rolling disc We demonstrate our approach on the vertical disc rolling in the plane which is a basic example in non-holonomic echanics, and we follow the conventions from [16,25]. The configuration space of the disc is R2 \u00d7 S1, where S1 denotes the unit circle, together with natural coordinates q = (x, y, \u03b8), where (x, y) is the position of the contact point P0 in the plane R2 and \u03b8 is the orientation of the disc in the plane, see Fig. 1. We assume that the radius of the disc is equal to 1 and that the rolling of the disc is with neither slipping nor sliding, i.e. the direct plane velocity of the disc is proportional to the angular velocity of the circular motion. In coordinates, the non-holonomic constraints are y\u0307 cos \u03b8 = x\u0307 sin \u03b8. (1) M a > > > o t P P L T i We can reformulate the system (1) as the Pfaff system dy cos \u03b8 \u2212 dx sin \u03b8 = 0, (2) and the annihilator of the system (2) forms the distribution D which is spanned by vector fields X1 = \u2202\u03b8 , X2 = cos \u03b8\u2202x + sin \u03b8\u2202y "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000066_012104-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000066_012104-Figure1-1.png",
        "caption": "Figure 1. The experimental setup",
        "texts": [
            " Upon contact of the metal anode with the liquid cathode, the combustion of the electric arc is initiated, then the metal electrode rises vertically to a distance of 5 mm from the surface of the electrolyte. During arc burning, the metal anode melts and the metal sprays under the action of plasma [6]. Liquid metal drops formation is observed, which quickly crystallize in the electrolyte. The goal is to study the parameters of the electric discharge during the metal powder formation. The gas discharge parameters were studied at the installation, the functional diagram of which is shown in Figure 1. It consists of an electrical power system 1, an electrolytic bath - 2, an electrode system - 3, an oscilloscope - 4, additional resistance - 5, a voltmeter - 6, an ammeter - 7, thermocouples - 8. Using the electrode system, the distance between the anode and the electrolyte solution was monitored. Using the oscilloscope 4, the shape of the applied voltage and current was controlled, and the voltage and discharge current were measured using a voltmeter and ammeter. Combustion of a gas discharge occurs between a cobalt-chrome alloy anode and an electrolytic cathode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002309_j.ijepes.2021.106860-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002309_j.ijepes.2021.106860-Figure3-1.png",
        "caption": "Fig. 3. (a) PMG cross-section perspective, (b) and (c) PMG rotor components.",
        "texts": [
            " The minimum required data necessary for the successful performing of magnetization are magnetic characteristics of PM, value of magnetic induction saturation and corresponding magnetic field value. Since neither manufacturer instructions or any technical data were available for the PMG, the missing magnet\u2019s characteristics had to be obtained by performing certain measurements, comparisons with magnets of similar materials and performing certain geometrical and electrical measurements on-site and further calculations. Cross-section perspective with rotor shaft and stator windings is presented in Fig. 3a. The mounting of rotor pole is given in Fig. 3b and c. It consists of housing and magnetic material. Main PMG parameters and dimensions are given in Table 1. PMG permanent magnets consist of 5 segments placed in an aluminium housing. PM material is AcAlCo 10,000 and its demagnetization characteristics and energy product are showed in the Fig. 4. The peculiarity of this material is a strong magnetic field of 160,000 A/m needed to achieve saturation flux density of 1.4 T, and value of reversible permeability of 5.5 \u00d7 10\u2212 6 Tm/A at 48,000 A/m. This suggests that relatively great number of ampere-turns must be provided for successful magnetization of permanent magnets of proposed characteristics (160 000 A/m, and saturation flux density 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001734_biorob49111.2020.9224410-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001734_biorob49111.2020.9224410-Figure4-1.png",
        "caption": "Fig. 4. Components of retractor-type knee joint.",
        "texts": [
            " Since the vehicle sensor lock is activated by the ratchet mechanism, even when knee flexion reproduced by pulling out the belt is prevented, knee extension reproduced by rewinding the belt can be done freely. In other words, when the vehicle sensor is activated, it is possible to prevent overflexion of the knee by stopping the belt from being pulled out, while simultaneously returning to the original posture by rewinding the belt. In addition, the push solenoid plays the role of the ball based on the law of inertia that activates the vehicle sensor. As shown in Fig. 4, by controlling the activation of the solenoid, we can prevent knee flexion at any time. In this way, we succeeded in developing a knee joint that can reproduce an arbitrary knee angle by controlling the on and off of the knee flexion prevention steplessly at an arbitrary timing, rather than adjusting the knee bending resistance using like a hydraulic cylinder in conventional robotic prosthesis. In this prosthesis, in order to measure the knee angle and the swing angular velocity of the lower limb, a rotary position sensor SV01A103AEA01 made by Murata Manufacturing Co"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003043_ssi52265.2021.9467023-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003043_ssi52265.2021.9467023-Figure2-1.png",
        "caption": "Fig. 2. CMUT technology and sensors. (a) Schematic of the unit cell. (b) Chip layout for a single channel sensor (1.4 x 1.6 mm\u00b2; CMUT aperture 1 mm). (c) Proximity sensors on TO-18. (d) Proximity sensor in housing and with interconnects.",
        "texts": [
            "00 \u00a92021 IEEE 2021 Smart Systems Integration (SSI) 20 21 S m ar t S ys te m s I nt eg ra tio n (S SI ) | 9 78 -1 -6 65 4- 40 92 -9 /2 0/ $3 1. 00 \u00a9 20 21 IE EE | D O Authorized licensed use limited to: University of Gothenburg. Downloaded on August 31,2021 at 11:10:56 UTC from IEEE Xplore. Restrictions apply. III. METHODS AND MATERIAL Fraunhofer IPMS designed and manufactured separate proximity and tactile devices based on its sacrificial layer technology for CMUTs, which is described elsewhere [20]. A schematic cross section of a typical individual sensor cell is depicted in (Fig. 2a). A single channel sensor chip (1.4 mm x 1.6 mm) comprises an array (1 mm diameter) of identical and electrically connected sensor cells (Fig. 2b). The proximity sensor chips were mounted on TO-18packages by chip bonding and wire bonding with GlobTop protecting the wires (Fig. 2c). Sensors are equipped with metal housing and electrical interconnects for system tests (Fig. 2d). A numerical simulation platform for air-coupled ultrasound based on the Elastodynamic Finite Integration Technique (EFIT, [21]) of Fraunhofer IKTS was developed and used to calculate the sound fields of the manufactured CMUT configurations and to investigate the interaction of the sound field with approaching objects. Characterization of the sensor devices on wafer level was performed by the means of capacitance-voltage tests and electrical impedance sweeps using a Keysight E4990A-120 analyzer in order to determine the operational voltage, the quasi-static capacitance and the resonance frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000786_s40430-020-2253-2-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000786_s40430-020-2253-2-Figure7-1.png",
        "caption": "Fig. 7 Comparison of the internal half-arc test and simulated standing wave wavelength in the grounding area under different speeds",
        "texts": [
            " At the same time, ABAQUS software is used to simulate the tire under the same speed range and boundary conditions to plot the deformation map of the tire. As the tire line speed is increased from 180 to 230\u00a0km/h, the angle of test and simulation is increased from 30\u00b0 to 45\u00b0. As the tire line speed increases, the standing wave gradually becomes apparent. As the angle increases, the standing wave wavelength of the tire sidewall becomes larger. The comparison of the standing wave waveforms on the sidewall of the tire in a semi-arc is shown in Fig.\u00a07 (images on the left are from experiments and on the right from simulation). The angle between the adjacent peaks of the sidewall is identified by the line passing through the center of the tire. It can be seen from Table\u00a02 that the error between the experimental value and the simulated value of the standing wave wavelength is small, and the maximum error between simulation and experimental results is less than 10%. As the speed increases, the wavelength also increases. At a maximum speed of about 230\u00a0km/h, the percentage error between the test and the simulation is 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure12-1.png",
        "caption": "Fig. 12. Decomposition of normal contact load.",
        "texts": [
            " (21), ' 1 1 ' 2 2 cos siny i z i zyi yzi x zxi yxi d d d d \u03b8 \u03b8 \u03b8 \u03b8 \u03b4 \u03c6 \u03b4 \u03c6 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u23a7 = + \u2212 + +\u23aa \u23a8 = + + +\u23aa\u23a9 . (21) Therefore, the curvature center distance after deformation ' 0A is, ' '2 ' 2 0 1 2A d d= + . (22) Then according to Eqs. (5) and (10), the normal contact load NiQ of the steel ball expressed by x\u03b4 , y\u03b4 , z\u03b4 , y\u03b8 , z\u03b8 can be obtained. To achieve the equilibrium equation of internal and external loads, xiQ \uff0c ziQ and yiQ need to be synthesized in five directions and balanced with the external loads in the corresponding directions. See Fig. 12. Eq. (23) is the equilibrium equations of external loads xF , yF , zF , yM , zM and ring displacement x\u03b4 , y\u03b4 , z\u03b4 , ,y z\u03b8 \u03b8 . For Eq. (23), it is a nonlinear equation set with five unknowns. The core iteration method we chose is the discrete Newton algorithm. Considering the particularity of the application. The following points need to be paid attention to: 1) Generally, choose x\u03b4 , y\u03b4 , z\u03b4 , y\u03b8 , z\u03b8 equal to 0 as the initial value, but when the clearance and the raceway curvature are relatively large, the solution efficiency is not high"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000048_j.biosystemseng.2019.10.007-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000048_j.biosystemseng.2019.10.007-Figure2-1.png",
        "caption": "Fig. 2 e Different arrays of arms of a MEA and examples of arms of each array.",
        "texts": [
            " Between these elements, some arrays of arms have been created. The arms rotate, as the MEA deforms in the axial direction. The displacement of the central disc with respect to the outer ring is performed thanks to the rotation of the arms. In Fig. 1 aMEAwith four rings, in addition to a central disc, is displayed. The number of arms in each array can vary, depending on the design. The smaller the number of arms in an array, the greater the potential displacement between the rings or ring and central disc could be reached. Figure 2 shows the different arrays of arms in a specific MEA and an example of an arm in each array. In the design of a MEA, it is possible to make many decisions in relation with the details of its geometry, as well as on its manufacturing process. For example, it is possible to define the number of arrays of arms and the number of arms in each array. In Fig. 2, it can be seen that the A-array presents 8 arms, while the B-array contains 7 arms. Moreover, C-array includes 5 arms, while there are 4 arms in the D-array. Additionally, the design of the arms can be performed with different shapes to favour or limit the rotation of the arms. The choice of thematerial of the MEA can also be a decision to be made, as well as the cutting technology applied for manufacturing the MEA. The thickness of the device is another geometrical parameter of the MEA that is analysed more in detail here"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003415_9781119352181.ch4-Figure4.21-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003415_9781119352181.ch4-Figure4.21-1.png",
        "caption": "Figure 4.21 End-winding configuration for form-wound coils.",
        "texts": [
            "121) The overall inductance is, from equations (4.38) and (4.39) Lew, circuit = P C Lew, phase belt (4.122) = (P C )( S 3P )(6C S )2 (kd1kp1Ns) 2 (Lew1 + Linternal) 2 (4.123) Thus the total end-winding leakage per phase can be expressed in terms of the total number of series-connected turns per phase Ns as Lew,phase = 6 S k2 d1k2 p1N2 s (Lew1 + Linternal) (4.124) The following formulas are taken from Liwschitz-Garik [5] which are the results partly of theoretical derivations and partly of experience. Refer to Figure 4.21 which illustrates an idealized portion of the end-winding region for the case of form-wound coils. Figure 4.21 shows two coils of a phase belt having q slots per pole per phase. Again, when the coils are now spread over q slots and then pitched Lew = q(kd1kp1ns) 2 pewlew (4.125) Here the inductance of equation (4.125) expresses the end-winding leakage for one entire phase belt comprised of q slots and q coils (two-layer winding). The corresponding inductance per slot is Lew\u2215slot = Lew\u2215q = k2 p1k2 d1n2 s pewlew (4.126) For machines of the type shown in Figure 4.21 (i.e., machines with form-wound stators), a suggested estimate of the specific permeance for the end winding is [5] pew = \ud835\udf070(1.2) (4.127) lew = 2(le2 + le1\u22152) (4.128) where the factor 2 takes into account the fact that the coil has sides projecting from either side of the machine, so that with le1 and le2 in inches and \ud835\udf070 = 3.192 \u00d7 10\u22128 henries per inch Lew\u2215slot = \ud835\udf070k2 p1k2 d1n2 s (2.4)[le2 + le1\u22152] (4.129) where in this result le1 and le2 are in inches and \ud835\udf070 = 3.192 \u00d7 10\u22128 henries per inch. The significance of le1 and le2 is seen from Figure 4.21. The method used in deriving equation (4.128) consists in treating the end-winding leakage as a revolving field in air. The presence of the iron is neglected in the derivation. Furthermore, it is assumed that the fluxes are confined to radial planes and are arcs of circles. Note that the pitch of the winding starts off at \ud835\udf0fp(ave) at the surface of the end lamination of the stack but drops off to zero by the time it reaches the end of the coil. Hence, an effective length of le1/2 is used in the formula. Clearly, the end-winding field is not truly a two-dimensional one but also spreads out axially. This effect was investigated experimentally by making approximate plots of the three-dimensional field. The effect of axial flux, neglected in the idealized analysis, is accounted for by the constant 2.4. For purposes of comparison, the permeance value of 2.4 occurs at a normalized overhang length of a = 2 when using the average value plotted in Figure 4.20. Note also the quantity te in Figure 4.21 which represents the air space between two insulated coil sides. In general, this quantity is controlled in order to provide insulation between adjacent coils. If te is fixed, it can be shown that le1 is related to te by le1 = p\ud835\udf0fp1(bc + te) 2 \u221a \ud835\udf0f2 s1 \u2212 (bc + te)2 (4.130) where \ud835\udf0fs1 and p\ud835\udf0fp1 are the slot and pole pitches measured at the middle rather than at the surface of the slot. The quantity bc is the breadth of the coil in the slot as opposed to the breadth of the slot itself. The value of le2 also depends upon the voltage and ranges from 0",
            "41), ns = 6CNs S The total end-winding leakage per phase can now be expressed in terms of the total number of series-connected turns per phase Ns as Lew = \ud835\udf070 S 3C2 ( 6CNs S )2 k2 p1k2 d1(2.4)(le2 + le1\u22152) (4.132) = 12\ud835\udf070 N2 s S k2 p1k2 d1(2.4)(le2 + le1\u22152) (4.133) A second alternative to calculating the end-winding leakage inductance of a formwound coil is to use an approximation based on the value calculated in Sections 4.8.3 and 4.8.4. Assuming that the area linked by the leakage flux is the same in both cases, by simple geometry it can be established that if the length a in Figure 4.18 is set equal to le2 + le1\u22152 in Figure 4.21 and b set to the pole pitch \ud835\udf0fp(ave); equation (4.124) can be used directly. While necessarily a crude approximation, the estimation is sufficiently good for an estimate. It should be mentioned that more accurate calculations of end-winding leakage inductance are available utilizing Neumann Integrals [3, 7]. The work involved, while intellectually satisfying, is rarely warranted given the uncertainty of the effect of eddy currents induced in the outer stator laminations. Figure 4.22 illustrates the end-winding region for a squirrel cage machine",
            "643 mH The mutual coupling between top and bottom coils for unity pitch is LlTB = 3N2 s le S1 pTB = 3(240)2 (8.67) 39.37 120 2.508 \u00d7 10\u22126 = 0.795 mH From Example 3 it was determined that the pitch of the stator coils is 0.8. From equation (4.90), the slot factor for mutual coupling is ksl = 3p \u2212 1 = (3)(0.8) \u2212 1 = 1.4 The total slot leakage inductance per phase is, from equation (4.89), therefore Llsl = LlT + LlB + ksl(p)LlM = 0.643 + 1.643 + (1.4)(0.795) = 3.400 mH Stator End-Winding Leakage Inductance From the additional information that has been given, the length le2 (Figure 4.21) is 1.25\u2032\u2032. The spacing between adjacent coil sides in the slot is specified as 0.0625\u2032\u2032. If this minimum spacing is maintained in the end-winding region, then the quantity te in Figure 4.21 is also 0.0625\u2032\u2032. Hence from equation (4.130), le1 = p\ud835\udf0fp1(bc + te) 2 \u221a \ud835\udf0f2 s1 \u2212 (bc + te)2 The pole pitch at the midpoint of the stator slot is \ud835\udf0fp1 = \u03c0 P (Dis + ds) = \u03c0 8 (24.08 + 2.2) = 10.32\u2032\u2032 The slot pitch at the midpoint of the stator slot is \ud835\udf0fs1 = P\ud835\udf0fp1 S1 = 8(10.32) 120 = 0.688\u2032\u2032 The length of the end-winding extension over the diagonal region is therefore le1 = (0.8)(10.32)(0.22 + 0.0625) 2 \u221a 0.6882 \u2212 (0.22 + 0.0625)2 = 1.859\u2032\u2032 The stator end-winding leakage inductance per phase is found from equation (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002752_10775463211010525-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002752_10775463211010525-Figure1-1.png",
        "caption": "Figure 1. Free-floating manipulator system Liang et al. (1998).",
        "texts": [
            " For matrices A,B2R n\u00d7n and x2R n, we have kABkF \u2264 kAkFkBkF (11a) jAxj \u2264 kAkF jxj (11b) The problem of mapping a FFM to a conventional fixed base manipulator which preserves both its dynamic and kinematic properties is discussed in Liang et al. (1998). This manipulator is called the DEM. Using this method for modeling a FFM not only simplifies dynamical and controller design analysis but also allows to build a conventional manipulator system for experimental studying of a space manipulator system without having to resort to complicated experimental setups to simulate the space environment. Figure 1 shows a n-link serial-chain rigid manipulator mounted on a free-floating base. The base is represented by link 1, and links of manipulator are numbered 2 to n + 1. Orientation of the space manipulator\u2019s base is represented by (\u03a6b 0, \u03b8b 0, \u03a8b 0) which are Euler angles. The space manipulator (i 1)th link and ith link are connected by Ji joint, \u03b8i is the rotation of Ji 0, and Ci 0, and C0 0 are the ith link\u2019s and total space manipulator\u2019s center of masses, respectively. \u03c1i 0 is the vector connecting C0 0 and Ci 0, the axis of rotation of Ji 0 is ui 0, Li 0 is the vector connecting Ji 0 and Ci 0, and also Ri 0 is the vector connecting Ci 0 and Ji+1 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000320_haz042-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000320_haz042-Figure16-1.png",
        "caption": "Figure 16. The maps \u03b2 and \u03b3 .",
        "texts": [
            " Consequently, f\u0302 n is isotopic to the identity on \u2202V and so the order of f is finite, which is a contradiction. Case 3: The case where both M0 \u222a Nbd(\u03c4 ) and M1 are solid tori. Set V0 := M1 \u222a X \u222a Nbd(\u03c4 ) (see again Fig. 15(ii)). Let D be the cocore of the 1-handle Nbd(\u03c4 ). The solid torus Cl(V0 \\ Nbd(D)) (\u223c= E(M0 \u222a Nbd(A))) is now a regular neighborhood of the trivial knot in S3. The disk D is then a primitive disk in V0 (see Cho (3, Lemma 2.1)). We can again think of f\u0302 as an automorphism of (S3, V0, D). By the argument in Lemma 5.5 of [4], f\u0302 2|V0 is a composition of \u03b1, \u03b2, \u03b3 shown in Fig. 16. Here \u03b1 comes from the hyperelliptic involution of \u2202V , \u03b2 is the half-twist of a handle and \u03b3 exchanges handles. Both \u03b1 and \u03b3 are order two elements, while the order of \u03b2 is infinite. We note that \u03b22|\u2202V0 is the Dehn twist along a separating simple closed curve c on \u2202V0. These elements satisfy \u03b1 \u25e6 \u03b2 = \u03b2 \u25e6 \u03b1, \u03b1 \u25e6 \u03b3 = \u03b3 \u25e6 \u03b1 and \u03b2 \u25e6 \u03b3 = \u03b1 \u25e6 \u03b3 \u25e6 \u03b2. Therefore, we can write f\u0302 2|V0 in the form f\u0302 2|V0 = \u03b1\u03b51 \u25e6 \u03b3 \u03b52 \u25e6 \u03b2n (\u03b51, \u03b52 \u2208 {0, 1}, n \u2208 Z). Thus, we have (f\u0302 2|\u2202V0 )4 = Tc 2n. Since c \u2229 \u2202A = \u2205, n should be 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002501_3447568.3448527-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002501_3447568.3448527-Figure1-1.png",
        "caption": "Figure 1: Map of potential fields around of an obstacle.",
        "texts": [],
        "surrounding_texts": [
            "Artificial neurons are a formal abstraction of the behavior of the biological [9] neuron. They form an ideal solution for certain problems that require reasoning, or that are of high complexity. This is due to their learning ability. A neuron is a nonlinear function, parameterized, with bounded value. A neuron contains 2 main elements: - The weights associated with neuron connections. - An activation function. The input values are multiplied by their corresponding weight and summed to obtain a sum Ui. \ud835\udc14\ud835\udc22 = \ud835\udc04(\ud835\udc31\ud835\udfcf, \u2026 . , \ud835\udc31\ud835\udc23, \u2026 . . \ud835\udc31\ud835\udc27) = \ud835\udeba \ud835\udc16\ud835\udc22\ud835\udc23 \ud835\udc31 \ud835\udc23 (1) Learning can be understood as a change in the capacity or behavior of an organism brought about by experience. [10] The learning algorithm will formulate the explicit rules that allow it to generalize. The formulation of the rules is done by the change of the synaptic weights which leads to the change of the behavior of the network; the change is carried out by a set of iteration which makes these networks able to react with new situations based on the experience passed. The neural understanding of knowledge is different; such knowledge can be in the following form: The optimal path is found by the synaptic weight vector W, hence a different coding of knowledge. This representation of knowledge is closer to the machine language, which makes the interpretation difficult for the human being in comparison with the other methods of representation of knowledge. For this reason, the networks of neurons are called black box. One of the goals of the research is to understand how knowledge is distributed within a neural network and how to extract that knowledge in a comprehensive way from a human being. The neural network approach is a method for modeling intelligence. These networks are able to solve problems in different areas. Their strong point lies in their learning ability. However the approach suffers from a major problem: the method of representing knowledge."
        ]
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure2-1.png",
        "caption": "Figure 2. The machining principle of the CZXB method.",
        "texts": [
            " Then considering the rotary motion of the worm blank, there are four feed movements totally. The four motions are controlled by CNC technology to work in the linkage mode. In theory, we can easily simulate the relative rotational motion between the worm blank and the cutting tool, as described in Section \u2018\u2018The traditional machining method\u2019\u2019. According to the Zhang and Li,16 a specific fourmotion linkage method, namely the CZXB method, is proposed. The processing principle of the CZXB method is shown in Figure 2. The machine coordinate system is a right-handed rectangular Cartesian coordinate system. The origin point of the machining coordinate system is O. The hourglass worm blank rotates around the axis O1-O1 at an angular velocity vector !1, and rotary movement is defined as C. The cutting tools are fixed on the rotary table. The rotary table rotates around the axis O0-O0 with the angular velocity vector !2, and the rotational movement is defined as B. The translation motion that rotary table translates along the radial direction of worm is defined as X",
            " Theoretical analysis of the CZXB method According to the generating theory of the planar enveloping hourglass worm, as shown in Figure 1, the generatrix S(2) rotates around the axis O2-O2 with the angular velocity vector !2, that rotary movement can be regarded as the planar motion of the rigid body. According to the basic theory of rigid body kinematics,18 the kinematics analysis of the CZXB method is shown in Figure 3. The fixed coordinate system g\u00bc [Og; xg, yg, zg] is the original position of the hypothetical tool gear. The position of the coordinate system g in Figure 2 is as follows: the axis xg is parallel to OZ and have the same direction, the axis yg and OX lie in the same line and have the opposite direction, and the axis zg and O2-O2 lie in the same line. The cutting tool is fixed on the rotary table. Generatrix S(2) is the main working surface of the cutting tool. Assume that the generatrix S(2) is a plane. S(2) and plane xgOgyg intersect at a straight line T1T2. Rotation axis of the rotary table and the plane xgOgyg intersects at the point O0. Any point P lies in the plane figures O0T1T2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000786_s40430-020-2253-2-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000786_s40430-020-2253-2-Figure1-1.png",
        "caption": "Fig. 1 Tire standing waves",
        "texts": [
            "ol.:(0123456789) Keywords High-speed camera\u00a0\u00b7 Radial tire\u00a0\u00b7 Standing wave\u00a0\u00b7 Cord force\u00a0\u00b7 Finite element When the rotational speed of the tire reaches a certain critical speed, the tire circumferential direction will show a static waveform relative to the observer following the center of the following axle, called a standing wave [1] (Fig.\u00a01). When the standing wave occurs, the rolling resistance of the tire will rise rapidly and the overall shape of the tire will no longer be round but wavy [2]. If the tire continues to run at this speed, (a) the temperature of the tire will rise rapidly in a short time, (b) the friction of the rubber molecules will resonate, (c) the circumferential direction of the tire and the distortion of the tire will increase, (d) the material properties will rapidly decrease and (e) the ply of the tread and the carcass will fall off, eventually resulting in a puncture"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001772_icmica48462.2020.9242909-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001772_icmica48462.2020.9242909-Figure3-1.png",
        "caption": "Fig. 3. Phasor diagram of DFOC [2]",
        "texts": [
            " But it requires complex mathematical modelling of the machine. The vector control requires instantaneous control of magnitude and angle of space vector. The dynamic modelling of the IM drive presented in Fig. 2. The measurement of the angle of space vector requires rotor position. The rotor position is measured with the help of sensors or by indirect technique. The sensor requires additional space and assembly of sensors, i.e. search coil or Hall Effect sensor or optical rotor position, which may degrade the dynamics of the drive [6]. The Fig. 3 illustrates the DFOC model. If the field angle is calculated from machine parameter and hall sensors, search coil, it is known as DFOC or if field angle acquired from machine parameter only then it is known as IFOC [2]. For the IFOC, the rotor flux-model reference adaptive system (RF-MRAS) is popular for low speed operation. The rotor position is estimated by reference and adaptive model. The reference model is derived from the stator voltage equation which is independent of speed to be measured",
            " 1 1 2 2 1 1 2 2 1 1 2 2 ls ms ms ms s ms ls ms ms ms ms ls ms L L L L L L L L L L L L L  + \u2212 \u2212     = \u2212 + \u2212     \u2212 \u2212 +    (4) qsi qri= \u2212ls s mL L L = \u2212lr r mL L L mL q s\u03c8 qr\u03c8q sv q rv sR rRe d s\u03c9 \u03c8 ( )\u2212e r dr\u03c9 \u03c9 \u03c8 dsi dri = \u2212ls s mL L L = \u2212lr r mL L L mL ds\u03c8 dr\u03c8dsv drv sR rRe qs\u03c9 \u03c8 ( )\u2212e r qr\u03c9 \u03c9 \u03c8 Fig. 2. d-q model of IM drive In the DFOC the rotor field angle is measured directly from the search coil or hall sensor and the rotor field angle is derived. The direct vector control scheme is well known and applied in high-performance AC drive. In this control scheme, the unit vector signals are generated in a closed-loop manner. The Fig. 3 presents the phasor diagram of direct vector control of IM. The conventional PI controller is used in speed regulation. From the phasor diagram, the direct control scheme can easily understand. The dqs axis is on the stator and dqr axis is fixed on the rotor. The rotor is moving at speed (\u0277), which is less than the synchronous speed. The synchronous rotating axis is ahead of the rotor axis by a positive slip angle \u03b8sl in case of motoring. From the slip relation, the rotor field angle is calculated directly by measuring actual rotor speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002037_j.engfailanal.2020.105126-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002037_j.engfailanal.2020.105126-Figure19-1.png",
        "caption": "Fig. 19. Upper body between arm 1 and 2.",
        "texts": [
            " 14, the suspension tower is modeled with finite element; each part of the tower has been built for the validation of the loads according to the section 3. In the Fig. 15, the material for retention tower has built, with the material influence for the calculations and constraints as elastic module, yield stress, steel type and components. Each component of the tower has been described in the Fig. 16 with the tower top. Fig. 17 for both upper arms, it applies to arm 1 and 2, Fig. 18 with the lowest arm of the tower, Fig. 19 in the upper body between arm 1 and 2, Fig. 20 for the body structure between arm 2 and 3, Fig. 21 and Fig. 22. Fig. 14. Tower model in finite element. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 The stage division of the ice disaster is composed of four stages: Preconditions (associated to the pre-failure period), steady-state progression and multiple failures (during disaster period) and restoration (post-failure period) [38]; recent papers have indicated effective resilience enhancement framework for towers installed, however, the recommendations for new assets are not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002923_s0263574721000588-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002923_s0263574721000588-Figure11-1.png",
        "caption": "Figure 11. Robots surrounding at t = 110 s with dynamically cluttered environment.",
        "texts": [
            " University of Glasgow Library, on 15 Aug 2021 at 00:27:22, subject to the Cambridge Core terms of use, available At around 75th iteration, algorithms with prediction and without prediction have pulled the robot toward the Passage_1. Algorithm without prediction chose this option because of observation 1. On the other hand, Algorithm with prediction chose this option because of the 2nd observation. At around 110th iteration (t = 110 s), states of Passage_1 and Passage_2 have changed, Key observations at the 110th iteration (t = 110 s): \u2022 Observation 3: Passage_1 has more free space compared to Passage_2 \u2022 Observation 4: Free space of passage_1 is decreasing, and free space of Passage_2 is increasing (Fig. 11) At the 110th iteration, algorithm without prediction has kept pulling the robot toward Passage_1 because of observation 3. However, algorithm with prediction chose to change the robot direction and at https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000588 Downloaded from https://www.cambridge.org/core. University of Glasgow Library, on 15 Aug 2021 at 00:27:22, subject to the Cambridge Core terms of use, available move toward the Passage_2, although it has less free space at that moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002260_cac51589.2020.9327221-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002260_cac51589.2020.9327221-Figure2-1.png",
        "caption": "Fig. 2. Kinematics Model",
        "texts": [
            " The influence of vehicle width can be eliminated by subtracting half vehicle width w from real road boundary, the kinematic constraint of the regional path tracking is simplified as: f \u2032 l (x) \u2264 y2 \u2264 f \u2032 r(x) (1) where y2 is the lateral position of the center of the mass of the semi-trailer, and f \u2032 l (x) and f \u2032 r(x) can be calculated as follows{ f \u2032 l (x) = fl(x)\u2212 w 2 f \u2032 r(x) = fr(x) + w 2 (2) III. VEHICLE MODEL In order to achieve better tracking performance of semitrailer, a novel simplified kinematics and dynamics models of tractor semi-trailer is established in this section. The schematic diagram of the tractor semi-trailer kinematics model is shown in Fig.2. Assuming that the semi-trailer is rigid body, according to the kinematics principle and the geometric relationship shown in Fig.2, the kinematic equation is as follows: \u23a7\u23a8\u23a9 x\u03072 = u2 cos(\u03c82 + \u03b22) y\u03072 = u2 sin(\u03c82 + \u03b22) \u03c8\u03072 = r2 (3) The meaning of each variable in the equation is shown in Table I. The curvature of the road is relatively small in most cases, and the yaw angle of the semi-trailer will be kept in a small range. Considering that the semi-trailer\u2019s sideslip angle \u03c8 is also small, the tractor semi-trailer kinematics relationship can be simplified as follows:{ sin(\u03c82 + \u03b22) \u2248 \u03c82 + \u03b22 cos(\u03c82 + \u03b22) \u2248 1 (4) By substituting Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000939_s12206-020-0434-7-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000939_s12206-020-0434-7-Figure7-1.png",
        "caption": "Fig. 7. Contact and deformation under load.",
        "texts": [
            " Because =e e wR f D , the distance between the center of curvature of the fixed outer raceway groove and the center of the ball can be given as: w w/ 2=( 0.5)= - -ej e el R D f D (1) Similarly, for the inner raceway groove: w( 0.5)= -ij il f D . (2) Further, the distance between the centers of curvature of the inner and outer raceway grooves can be estimated using the following equation: w( 1) .= + -i eA f f D (3) According to the relationship between the load along the virtual bar direction ejQ and the ijQ elongation of the virtual bar system after loading in Fig. 7, the formulas for calculating the elongation D ejl and D ijl of the virtual bar system are as follows: 2 3\u00e6 \u00f6 D = \u00e7 \u00f7\u00e7 \u00f7 \u00e8 \u00f8 ej ej ej Q l K (4) 2 3\u00e6 \u00f6 D = \u00e7 \u00f7\u00e7 \u00f7 \u00e8 \u00f8 ij ij ij Q l K (5) In the formula: ejQ and ,ijQ respectively represent the contact loads of the ball with outer and inner rings ejK and ,ijK respectively, represent the deformation-load coefficients of ball and inner and outer rings. According to the relationship between the radius center of curvature difference of outer ring and inner ring raceway and the position of ball center of high-speed ball bearing after loading in Fig. 7, the calculation formulas of the final length ' ell and ' ill of virtual bar system are determined as follows: ' w( 0.5)= + D = - + Del ej ej e ejl l l f D l (6) ' w( 0.5)= + D = - + Dil ij ij i ijl l l f D l (7) In the formula: w=e ef D R , w=i if D R ; ' ell and ' ,ill respectively, indicate the distance between the radius center of curvature difference of outer race and inner race raceway and the position of ball center when the bearing is not loaded. wD denotes the diameter of the ball. ef and ,if respectively, indicate the curvature coefficients of the outer and inner ring channels, eR and ,iR respectively, indicate the radius of curvature of the outer and inner ring channels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001872_s11424-020-9144-6-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001872_s11424-020-9144-6-Figure6-1.png",
        "caption": "Figure 6 Experimental set up: One DOF telerobotic system with haptic",
        "texts": [
            " Remark According to the proposed control method, for all j = m, s, rj = q\u0307j + \u03bbqj is also bounded and is able to converge to zero under ideal network conditions. In this case, after the human force is removed, the master and slave robots will move to zero position, i.e., qm = 0 and qs = 0. In practice, however, it is expected in this scene that the master and slave robots achieve position synchronization, i.e., qm \u2212 qs = 0. The system used for the experiment consists of two single DOF robots operating as the master and slave systems as shown in Figure 6. The main components of the system are actuators, drivers, sensors, data acquisition cards and the communication channels. Two DC motors are used as the actuators with nominal voltage and power of 22V and 40W, respectively. One of them is located on the master side and associated with an operator and the next one is located on the slave side and interacts with the environment. For driving these motors, two analogue drivers is applied. To read motors positions, an optical encoder with the resolution of 400 pulse per turn is used",
            " To access the information regarding to the force applied to the master and the slave, two high precision load cells are employed. In this project, because of the hardware limitation, two computers as the target and the host are employed. In fact, the target computer is used for data gathering via some data acquisition cards and the host computer is used for implementing the control algorithm. For communicating between this two computers, UDP protocol is also employed. Also, the parameters of the designed controller in the experiments is set to Ks = Km = 1.2. The general view of this system is illustrated in Figure 6 (b). First, the performance of the conventional wave variables adopted in [23] is first provided. Then, the results corresponding to the proposed control structure are given. Figure 7 depicts the position and force tracking in the pure channel. In this experiment, the initial position of the slave is 25 degrees far from the environment. It can be seen that during the time that the slave is in the free space (e.g., from 0 to 6 s, 8 s to 10 s and etc), the slave perfectly follows the master. However, during the time that the slave robot is in contact with the environment (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000503_012020-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000503_012020-Figure1-1.png",
        "caption": "Figure 1. Design and dimensions of a waterwheel simulation area",
        "texts": [
            " Whereas \u03c9isangular speed: \u03c9= 4vt / [D(1+ Cos\u03b8/Sin \u03b1)] (13) n med rotation speed: n med = (30/\u03c0n) x S \u03c9j (14) 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 \u03b8 = cos-1[{(0.5 x D) \u2013 (B x h)}/ (0.5 x D)] (15) \u03b1 = (90 \u2013 \u03b8 ) + \u03b8 x (m / m+1) (16) m is a constant of 1\u2264m\u22646 To find out the Radial speed, the equation below is used: Vr= v( Vam2 \u2013 2 x Vam x Vt x sin a + Vt2) (17) Tangential velocity use: Vt = (Vam x sin \u03b1) x 0.5 x [1+( cos\u03b8/sin \u03b1)] (18) tan \u03b3 = (Vt x cosa) / (Vam - Vt x sin a) (19) The design and dimensions in Figure 1. above are obtained from the calculation of 1 to 15. In accordance with the purpose of this study, a waterwheel planning was made which was later simulated using CFD. CFD is a method of numerical fluid flow approach with the help of a computer, which can work in low water flows. This condition is suitable with the character of the river flow in the regions therefore it can be applied in Indonesia with many small islands. The vertical waterwheel design and construction obtained as shown in Figure 1, dimensions of blade length 600 mm and width of blade 100 mm with number of blades 12 pieces, to determine the speed of water flow before and after passing through the waterwheel. Waterwheel modelling using Ansys Student Version 19.1 is calculated based on flow assumptions, with an initial flow rate of 4.85 m/s by entering the gravity acceleration value 9.81 m/s2. One of the models used is k-epsilon. Modelling by using this system provides advantages in the form of computational power efficiency, stability of numerical calculations and accuracy of the resulting solution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003101_s42417-021-00349-z-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003101_s42417-021-00349-z-Figure3-1.png",
        "caption": "Fig. 3 Co-ordinate system of Hybrid Gas Foil Bearing",
        "texts": [
            "\u00a0(20), Pa is the atmospheric pressure, P\u0304 is the normalized pressure, H is the film thickness, H\u0304 is the normalized film thickness, Z\u0304 is non-dimensional axial bearing length, \u039b is the bearing number, is the angular velocity of the journal, is the non-dimensional time,C is the clearance and R is the radius of the bearing. In case of gas foil bearing, the foil deflection has to be considered in the fluid film thickness equation. Therefore, the film thickness equation for the geometry of the bearing described in Fig.\u00a03, can be written as, (21)H\u0304 = 1 + e\u0304x cos \ud835\udf03 + e\u0304y sin \ud835\udf03 + w\u0304t where,w\u0304t is the non-dimensional elastic deformation of the foil structure, e\u0304x and e\u0304y is the non-dimensional displacement along X axis and Y axis respectively. In the current investigation, the foil structure has been formulated based on Classical Plate Theory (CPT) using the methodology given in Basumatary [19]. The deformation depends on the bump foil stiffness, Kf and the average pressure across the bearing width. Thus, the non-dimensional deflection of the bump is given by, 1 3 where, S = Pa Kf C is the compliance coefficient, D\ufffd 11 = C PaR 4 D11 , D\ufffd 12 = C PaR 2l2 D12 , D\ufffd 22 = C Pal 4 D22 , D\ufffd 66 = C PaR 2l2 D66 and D11 = D22 = Et3 t 12(1\u2212 2) ;D12 = Et3 t 12(1\u2212 2) ; D66 = Gt3 t 12 in which, is the Poisson\u2019s ratio, E is the Young\u2019s modulus of the top foil, tt is the thickness of the top foil and G is the shear modulus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001656_etfa46521.2020.9211891-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001656_etfa46521.2020.9211891-Figure7-1.png",
        "caption": "Fig. 7: Search movement cases",
        "texts": [
            " In this context, an interesting concept that does not require direct force/torque measurement is the robotic peg-in-the-hole assembly strategy presented in [14]. A superposition of a linear movement in x- and a sinusoidal movement in y- direction of the end effector frame FE allows the robot to search for the entry and to handle the clearance between connector and mate. Therefore, it is suitable to represent the y- movement as a function of the x- movement. It is assumed that the position of the connector\u2018s mate is known with an accuracy of \u00b13 mm in x- and z- as well as \u00b12 mm in y- direction. Fig. 7 shows four search movement cases with displacement \u2206x \u2208 R of the x-axis and amplitude a \u2208 R of the sinusoidal movement of the y-axis. These variables are used for the calculation of the movement and the four cases just differ in the sign of \u2206x and a. This allows a variation of the search process and is necessary if the connector moved aside of its mate, as shown in Fig. 6, and therefore requires another retry. The red area of Fig. 7 represents the jack, which is rigidly mounted on the contact flange and the black part shows the moving connector. The blue arrow illustrates the executed search movement. The linear movement is given by x(t) = x(\u03c3(t)) = x0 + \u03c3(t) \u2206x (1) with start position x0 \u2208 R and displacement \u2206x, which is basically just the difference between start and end position of the movement. The parameter \u03c3(t) \u2208 [0, 1] removes the direct time dependency of x and allows the scaling of the path. The corresponding time derivatives are x\u0307(t) = x\u2032(\u03c3(t)) \u03c3\u0307(t) = \u2206x \u03c3\u0307(t) (2) 1467 Authorized licensed use limited to: Auckland University of Technology",
            " In some cases, the plug might already be inside its connector and so this needs to be checked. Therefore, the robot tries to apply a force of \u00b14 N in x- and y-axes one after another. If this is possible for both directions without exceeding defined position limits in the search plane (x/y-plane) or detecting a force jump in the z-direction, then the plug is really inside the jack. Fig. 8 shows such an optimal force path. The sign of the force in the search plane equals the sign of \u2206x and depends on the current search case (see Fig. 7). If the forces in x and y cannot be reached, then the plug is definitely outside the jack and it is necessary to start the sinusoidal search movement. Fig. 9 shows the case where the force in z jumps during a started search. This indicates a potential entry of the jack and requires a subsequent check via applying forces in both directions of the search plane like 1468 Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 02,2020 at 11:44:48 UTC from IEEE Xplore. Restrictions apply. explained before. The plug is inside its connector in the case of a successful check procedure. In the cases of another force jump in z or a robot movement beyond predefined position limits, then the plug is outside the jack. Consequently, it is required to move the robot back to the start without causing any collision and to vary the search movement as shown in Fig. 7. This repeats until the entry of the jack is found. Fig. 10 represents the optimal force path where another force jump in z happens once before the check procedure could be started. Then the entry of the jack could be found via the second search movement case. For a safe contact between a robot and its largely unknown environment, mostly force control is used. The robot should follow a precalculated trajectory while also develop certain forces in several axis. This is possible via a parallel force and position control [15], [16], which is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000703_s12206-020-0122-7-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000703_s12206-020-0122-7-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of complex structures.",
        "texts": [
            " The detailed model using FEA was verified by comparing with the experimental results and is for verification of boundary condition. Among the three equivalent models, an equivalent model which is similar to analysis results of the detailed model, was selected and was applied to the wheels of the commercial vehicle to verify the equivalent model. Since it is difficult to perform bolt fastening and loosening tests on the wheel structure of a commercial vehicle, we propose a complex structure to replace it. The complex structures consist of bolts, nuts and three fastening component, as shown in Fig. 2, which is the structures similar to the wheel of a commercial vehicle. The components of the complex structures are shown as Fig. 3. The bolt model of the complex structures is composed of a M12, the clamped part has a width of 100 mm, a width of 150 mm, and thicknesses of 6 mm and 15 mm, respectively. The thickness is the same as the wheel and hubdrum structure of a commercial vehicle. The M12 bolts are high-strength general hexagon bolts with heat treatment and zinc-coating produced according to hexagon bolt with large width across flats for high tensile structural bolting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000389_phm-qingdao46334.2019.8942977-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000389_phm-qingdao46334.2019.8942977-Figure2-1.png",
        "caption": "Figure 2 Test rig of inner race fault bearing: (a) the test rig, (b)the rolling bearing with the fault in inner race.",
        "texts": [
            " ( )log ,t ts a\u03b8\u03c0 is used to transform the policy into probability; then take the partial derivative with respect to the probability to calculate the gradient; tv is calculated by using policy gradient theorem and the stochastic gradient ascent algorithm. In the third step, each time an action is generated, build a corresponding bandpass filter to filter the original vibration signal; next, ues the Hilbert transform to the filtered signal and calculate the kurtosis according to the envelope spectrum of the filtered signal. The reward of each action can be set as ,l hESk . The reinforcement learning system will converge after iterating a certain number of times, and the ,l hESk would be the maximum. Figure 2 (a) shows the test rig, and the real data was collected from the bearing with an inner race fault (shown in figure 2 (b)). The data is applyed to validate that the proposed method is effective. The bearing type is 6205-2RS SKF with the seeded fault on the inner race. The fault is generated by using wire-electrode cutting with wire width equaling to 0.2 mm. During the data collection, the speed of the motor kept at 1046 rpm (revolutionper-minute). A force of 1 kN from the mechanical loading structure was loaded on the bearing. The NI acquisition unit (NI PXI-4462) was used, and an accelerometer was installed on the bearing pedestal to gather data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002910_s12206-021-0512-5-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002910_s12206-021-0512-5-Figure2-1.png",
        "caption": "Fig. 2. ADAMS high-precision dynamic model.",
        "texts": [
            "\" The friction coefficient of joint surface is obtained by measuring different materials and surface treatment. Table 1 shows the results. Apply ADAMS/Car to obtain a high-precision dynamic model by setting up dynamic parameters (geometry), quality parameters (mass, centroid and moment of inertia), mechanical parameters (spring, damping, etc.), and external parameters (road conditions). Suspension assembly, rear suspension assembly, body model, and steering model are integrated to form a complete vehicle dynamics model as shown in Fig. 2. Load calibration test and road test simulation are used to calibrate the actual working load, so as to ensure the accuracy of bolts under various actual working conditions (such as forward, rear, emergency braking, over-deep pit, steering braking, driving, etc.). First, strain gauges are attached to the bolted joints to collect signals (see Fig. 1(a)). Second, load calibration test (see Fig. 1(b)) and road load spectrum measurement test (see Fig. 1(c)) are used to obtain test data to verify the simulation model and to finally generate load conditions (an example shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000412_phm-qingdao46334.2019.8942851-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000412_phm-qingdao46334.2019.8942851-Figure7-1.png",
        "caption": "Fig. 7. Schematic of apparatus to electrically characterize the printed electrospray devices.",
        "texts": [
            ", the arrays of emitters are hexagonally packed with 888 \u00b1 7 \u03bcm emitter pitch); in these devices each emitter spout is fed by a 12 mm long tapered internal channel with 400 \u00b1 15 \u03bcm diameter at the emitter spout and 498 \u00b1 10 \u03bcm diameter at the emitter inlet (Figure 5). In addition, devices with as many as 236 straight electrospray emitters in 1 cm2 (i.e., the arrays of emitters are hexagonally packed with 700 \u00b1 7 \u03bcm emitter pitch) were made; in these devices each emitter spout is fed by a 12 mm long internal channel with 377 \u00b1 15 \u03bcm diameter (Figure 6). An schematic of the apparatus used to characterize the devices is shown in Figure 7. The electrospray array is clamped to an aluminum chuck using a set of polymer screws. The chuck has a feedthrough that supplies liquid to the device, and an O-ring is used to seal the interface between the device and the chuck. The liquid is delivered to the chuck using a Harvard 33 syringe pump. The same screws that clamp the device to the chuck are used to integrate the extractor electrode, i.e., a laser-cut 250 \u03bcm-thick stainless steel plate with apertures that line up with the emitters of the array, and the collector electrode, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure38.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure38.3-1.png",
        "caption": "Fig. 38.3 a Center of mass of the frame Gf and b common center of mass Gsf coordinate as measured from the pivot of point c",
        "texts": [
            " The general equation for estimating the center of mass was adopted for the final calculation. The center of mass for the rotating part of the frame can be computed after defining its geometry as well as material properties of the frame which is used as an input for computing the same in CAD software such as Catia V5 which is adopted for current analysis. Thus, the simulated result for the center of gravity ormore commonly the center of mass of the frame is located near to the saddle as shown in Fig. 38.3a; the coordinates are (Xf, = 204.12mm and Yf = 400.84mm) and denoted byGf. Although the present analysis is undertaken with a human\u2013machine context, a single or common center of mass is mandatory in order to handle or apply an equilibrium equation; thus, to determine the common center of mass between the frame Gf and the subject Gs, the following combined equation is used. ( Xs f,Ys f ) = [ msxs + m f x f ms + m f , ms ys + m f y f ms + m f ] (38.1) where Xs f,Ys f are the common center of mass coordinates from pivot of point c; ms,m f are the masses of the subject and the frame, respectively; similarly, xs, x f ys and y f are centroidal distance from the center of mass of the subject and the frame with respect to X and Y axis, respectively. The common center of mass is located slightly below the center ofmass of the subject Gs as shown in Fig. 38.3b and denoted as Gsf. Figure 38.3b comprises a geometrical parameters such as the actuator inclination angle measured counter clockwise from horizontal and the main frame inclination angle \u2205 measured clockwise from horizontal. The worst case of transfer phase is considered for this analysis, which requires the maximum effort from the actuators, and this occurs during the initial phase of lifting the subject; thus, for the present subject\u2014device interaction, it occurs when the main frame is at an angle of \u2205 = 60 and from the horizontal and when the actuator angle is at = 65.51\u00b0 from horizontal as shown in Fig. 38.3b. In order to estimate the maximum actuator force (Fa) needed for the worst-case scenario, analysis of stability equation will sound more reliable, and it was adopted to precisely compute the optimum actuator load Fa; the stability equation states that stabilizing moment MS \u2265 MD destabilizing moment; thus, the actuator load Fa with respect to the combined mass of frame and subject (msf ) can be estimated from equation of stability as: Fa \u2265 msf \u00d7 g \u00d7 Xsf ACY cos + ACX sin (38.2) The final decisions for locating the effective position of the actuator are one of the crucial design criteria in the process of a linear actuator design for patient lifting mechanism; therefore, based on the orientation of a load application coordinate point A (ACX, ACY), the corresponding force Fa which is expected from the actuator will change radically due to a change in the coordinate point (A) for the same subject in a given posture as discussed previously"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000547_iros40897.2019.8967922-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000547_iros40897.2019.8967922-Figure6-1.png",
        "caption": "Fig. 6: Configuration space of an aerial manipulator (i.e., a multi-rotor with a 2-Dof arm) in our set-ups.",
        "texts": [
            " 5a\u20135b and Figs. 5c\u20135d show the simulation results from each case, respectively. The waypoint spaces are shown in yellow. In Fig. 5a and Fig. 5c, the trees from informedRRT* are shown in bidirectional manner. The blue and green lines show the edges in T and T I , respectively. The red circles show the sampling spaces. In Figs. 5b and 5d, each red line shows the resulting trajectory, respectively. We consider an aerial manipulator which is a combined system of a multi-rotor and 2-DoF robotic arm (see Fig. 6). The state vector of the combined platform is represented by the collection of state variables of the multi-rotor and robotic arm as q = [p>b \u03a8 \u03b7>]>. The vectors pb = [x y z]>, \u03a8= [\u03c6 \u03b8 \u03c8 ]>, and \u03b7 = [\u03b71 \u03b72] > denote the body position of multirotor, Euler and joint angles of multirotor and robotic arm, respectively. As described before, we assume that we can pick an object in any direction and give the location of an object as a constraint. Thus, the waypoint for pick-andplace task is given for the end effector position and equation (1) can be formulated as, fw(x,y,z,\u03c6 ,\u03b8 ,\u03c8,\u03b71,\u03b72) = p\u0304e\u2212 p\u0304w (4) = g2(\u03b72)g1(\u03b71)g0(\u03a8)p\u03040\u2212 p\u0304w (5) = 0, where gi\u2019s for i = 0,1,2 are the transformation matrices in [15]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002903_j.mechatronics.2021.102593-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002903_j.mechatronics.2021.102593-Figure3-1.png",
        "caption": "Fig. 3. Sensing system with loadcell to measure mass distribution.",
        "texts": [
            " As a result, the cockpit orientation can be estimated from both IMU and optical sensors minimizing accumulation error and maximizing the robustness of the optical sensor. The SMP in Fig. 1 is designed with the assumption that PG (CG) is aligned with P0 (CR). However, it is impossible to satisfy due to weight of an operator and equipment. PG should be estimated to calculate G in (13) and apply it into the control system. The unbalanced weight caused by the mismatch between PG and P0 incurs overloads for particular actuators, accelerating slipping motion. The loadcell sensor is implemented to measure transferred mi from the compression spring, as shown in Fig. 3a, where a cover protects the sensor in Fig. 3b. According to the cockpit rotation, in (20), PG is estimated by the changes in the loadcells, mi (i=1,2,3,4). [ M1 M2 ] = [ C\u03b8C\u03c8 \u2212 C\u03b8S\u03c8 S\u03b8 C\u03d5S\u03c8 + S\u03d5S\u03b8C\u03c8 C\u03d5C\u03c8 \u2212 S\u03d5S\u03b8S\u03c8 \u2212 S\u03d5C\u03b8 ] PG (20) Mi = miqi \u2212 mi+2qi+2 mi + mi+2 , i = 1, 2 (20a) where S and C indicate sine and cosine functions of the subscript angle, respectively; PG = [PG,X PG,Y PG,Z]T. In (20a), Mi represents the weight distribution of the cockpit sphere measured from the loadcells, and it is computed by weight from the loadcells and the position of the stages"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure44.9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure44.9-1.png",
        "caption": "Fig. 44.9 Glame coupled with both small and large dia. plates",
        "texts": [
            "H er e, th e G la m e is in cl in ed at \u22124 \u00b0 to m ai nt ai n st ab ili ty Pl at e w ith an ex te nd ed co lla r. H er e, th e G la m e is de cl in es at 32 \u00b0 to m ai nt ai n st ab ili ty Dishware plates are segregated in mainly 2 sizes, i.e., food plate which is big and dessert plate which is small. The experiment shows how Glame interacts with the plates of different sizes. Large or small, Glame braces both the plates impeccably. The claw lock is designed in such a way that it barely hinders the good area of the food plate (Fig. 44.9). In the left, food plate generally ranges from 11 to 13 in. in diameter. Glame clamps the plate perfectly in spite of utilizing very little area on the front. In the right, dessert plate generally ranges from 7 to 9 in. in diameter. The thickness of the claw lock is so absolute that it braces both the size of plates without any imbalance. A pictorial glance of \u201cGlame\u201d in-use. It conveys the confidence to clasp singlehandedly. Also, it does not hinder the dishware well which furnishes the sense of hygiene while consuming food"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001497_kem.861.113-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001497_kem.861.113-Figure5-1.png",
        "caption": "Fig. 5 Heating stage of plasma spraying process",
        "texts": [
            " Plasma spraying technology is also a very promising repair technology, which is widely used in the wear protection of engine thermal barrier coating system. T. Kalfhaus et al. [10] from Germany studied the influence of plasma spraying process parameters on the microstructure. The results showed that: the increase of substrate temperature, the decrease of powder feeding speed and viscosity will lead to the decrease of porosity; while when the solidification speed is slow, there will be smaller pores; similarly, the hot isostatic pressure treatment of the material after plasma spraying will also lead to the closure of pores (Fig. 5). Ye et al. [9] of the State Key Laboratory of tribology, Tsinghua University, studied the high temperature alloy after brazing repair. It was found that there were a large number of needle and block borides in the diffusion affected area after brazing repair. After heat treatment at 1180 \u2103/1190 \u2103, most of the borides dissolved, and the material structure and tensile properties in the welding area were improved. Although the superalloy has good thermal stability and high thermal strength, it often has defects such as crack, burn and tissue degradation due to the influence of many factors in the application process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003988_09544070211035859-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003988_09544070211035859-Figure7-1.png",
        "caption": "Figure 7. Brush model considering the carcass elasticity.",
        "texts": [
            " Analysis of small side slip angle characteristics of brush model considering tire wear The main mechanical properties of the tire in the small side angle are the stiffness characteristics, which represent the tread bristles distribution stiffness of the brush model. The brush model is a simplified tire physical model,24 which can reflect the basic mechanical properties of the tire. A tire brush model considering the elasticity of the carcass25 is established in this paper. The schematic diagram of model is shown in Figure 7 and its cornering properties expression is as follows. Dy(x)= yt(x) yc(x) yt(x)= yc0 + a u (a x) tga yc(x)= yc0 + yu(x)+ ycb(x) 8>< >: \u00f011\u00de Dy(x)= (a x) Mz Nu tga Fy Kcb j x a \u00f012\u00de where: j is a general bending deformation function of the carcass, j(x)= x2 1 2 . Nu is the torsional stiffness of the carcass. Therefore, the lateral force and the aligning moment can be expressed as Fy = \u00f0a a ktyDy(x) dx Mz = \u00f0a a ktyDy(x) x dx 8>< >>: \u00f013\u00de where e0 =Kty Kc0 =2akty Kc0 eb =Kty Kcb=2akty Kcb eu =Nta Nu = 2 3 a3kty Nu 8>>< >>: \u00f014\u00de In equation (14), Kty is the lateral stiffness of the tread, Kc0 is the lateral translational stiffness of the carcass, Kcb is the lateral bending stiffness of the carcass in the footprint, and Nu is the torsional stiffness of the carcass in the footprint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001153_01691864.2020.1782260-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001153_01691864.2020.1782260-Figure12-1.png",
        "caption": "Figure 12. Demonstration key poses inserted (left in each subfigure) and the collision-free motion paths (right in each subfigure) generated with (a) Policy 1 and (b) Policy 2 of L insertion. Both use the planning result in first row under each policy.",
        "texts": [
            " Although the single planning time increased a little bit resulting from fewer key poses. Demonstration key poses inserted Table 1. Motion planning result of L insertion. Selection policy Inserted pose number Single planning time (s) Total time (s) Policy 1 11 1.333 22.086 11 0.971 20.060 11 1.013 20.572 11 0.932 21.795 11 1.156 22.498 Policy 2 1 1.526 1.531 4 1.403 6.285 5 1.809 8.704 5 1.647 9.398 3 2.961 6.031 and the eventual obtained collision-free motion paths of L insertion generated with Policy 1 and Policy 2 are shown in Figure 12. Table 2 shows the results of motion planning of tenon insertion. Similar to L insertion, Policy 2 also involved fewer demonstration poses and cost relatively less total time than Policy 1. Meanwhile, the single planning time increased a little bit resulting from fewer key poses. Demonstration key poses inserted and the eventual obtained collision-free motion paths of tenon insertion generated with Policy 1 and Policy 2 are shown in Figure 13. Table 3 shows the results of motion planning of peg-rot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000144_elecsym.2019.8901645-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000144_elecsym.2019.8901645-Figure2-1.png",
        "caption": "Fig. 2. (A) mobile robot front view, (B) mobile robot top view, (C) mobile robot base isometric view",
        "texts": [
            " For the localization and mapping purpose, mobile robot in this research is using a LRF sensor. Therefore, LRF sensor on mobile robot must be put in a place with a height around 20 cm above the ground. By placing LRF sensor 20 cm above the ground, mobile robot can capture all objects that could be an obstacle to the course of the mobile robot. The real mobile robot that will be builded is using an omni directional wheel. To make mobile robot design as similar as real robot in real world, the mobile robot base design is equipped with three omni directional wheels. Fig. 2 shows mechanical design of mobile robot base. (A) (B) III. LOCALIZATION AND MAPPING STRATEGY The goal of simultaneous localization and mapping (SLAM) is to know the position of mobile robot and to build a map based on the environment explored by mobile robot. To obtain drawings of the environment explored by mobile robot, LRF sensor is used in this research. LRF sensor acquire surroundings data by measuring distance between LRF sensor position and object or obstacle around mobile robot. Hokuyo-URG-04LX is a type of LRF sensor provided by V-rep simulator that has a 270o scanning area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002312_indicon49873.2020.9342069-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002312_indicon49873.2020.9342069-Figure1-1.png",
        "caption": "Fig. 1. Showing 6 degree of freedom positions, velocities, forces and moments of ship. [46].",
        "texts": [
            " M v + C \u03bd \u03bd + D \u03bd \u03bd + g \u03b7 + g = \u03c4 + \u03c4 + \u03c4 1 \u03b7 = J \u03b7 \u03bd 2 where M is 6 \u00d7 6 mass matrix represents sum of rigid body mass and added masses, C represents 6 \u00d7 6 Coriolis matrix due to rigid body and added mass, D represents 6 \u00d7 6 damping matrix which is combination of linear and non-linear damping, g \u03b7 vector represents 6 \u00d7 1 hydrostatic matrix and g is a matrix of control inputs and represents the thruster forces which are combination of forces and moments causing motion, is a position vector represents combination of translational and angular positions, is a velocity vector combination of both angular and positional velocities [46] is presented in fig1 Modelling of AUV involves rigid body dynamics along with hydrodynamic modelling. The hydrodynamic modelling of AUV comprises of added masses, hydrodynamic damping, Coriolis due to added mass is proposed in literature datcom [26]. Datcom is a data compendium for obtaining aircraft stability properties. It can be easily applied to AUVs by considering non- dimensional coefficients. The hydrodynamics of AUV is expressed in [3, 6]. The added masses of vehicle is due to movement of vehicle in presence of water medium, damping matrix consists of linear and non-linear damping due to cortex shedding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000555_j.jmapro.2020.01.054-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000555_j.jmapro.2020.01.054-Figure19-1.png",
        "caption": "Fig. 19. Normal residual stress distribution of the components fabricated by AM based on differently shaped substrates in (a) GR-3, (b) PL, and (c) BO-3 schemes.",
        "texts": [
            " In addition, the high transverse tensile stress zone is mainly concentrated below the interface between the AM part and the substrate, and its range is small. Because the interface positions in the models are different, the heights of their high transverse tensile stress zones are also different. For example, the high transverse tensile stress in the model in GR-3 mainly concentrates below the groove, while the stress in BO-3 is concentrated below the bonding zone between the boss and AM part. The normal stress distributions of the models in GR-3, PL, and BO-3 are shown in Fig. 19. From this figure, it can be observed that there is a normal compressive stress zone in the middle of the three models. It is well known that a certain amount of compressive stress is advantageous for the structural integrity of a component. However, the normal stress distribution of the model in BO-3 is different from that of the other two models. Its compressive stress zone is not around the bonding zone between the substrate and AM part, but appears below this zone, whose height is close to this region in PL"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002350_iros45743.2020.9341242-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002350_iros45743.2020.9341242-Figure2-1.png",
        "caption": "Fig. 2. Mechanism of developed EPW (IMU, inertial measurement unit) [12]",
        "texts": [
            "00 \u00a92020 IEEE 4114 20 20 IE EE /R SJ In te rn at io na l C on fe re nc e on In te lli ge nt R ob ot s a nd S ys te m s ( IR O S) | 97 8- 1- 72 81 -6 21 2- 6/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ IR O S4 57 43 .2 02 Authorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 16:29:29 UTC from IEEE Xplore. Restrictions apply. a slider. In addition, we evaluated the stability while going down stairs and the effectiveness of the proposed control method. Front and side views of the mechanism proposed by Shino et al. are depicted in Fig. 2, and its specifications are listed in Table I. The proposed mechanism comprises two rotary links between the front and rear wheels on each side, and a slider that moves the seat back and forth. The user is assumed to be fixed to the seat with a seat belt. As it has been experimentally confirmed that the behavior of this device has low effect on the user, a payload can substitute the real user [12]. The device has a slider actuator under the seat and rotary links and wheel-drive motors. Shino et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000032_chicc.2019.8866066-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000032_chicc.2019.8866066-Figure1-1.png",
        "caption": "Fig. 1: A Seven-degree-of-freedom vehicle dynamics model",
        "texts": [
            "Section 3 introduces the parameter estimation algorithm.The paper ends with conclusion in Section 4. For the vehicle longitudinal and lateral motion control system design, a control model is required with steering angle as inputs and vehicle states as output.In this section, the description of control model is divided into three parts. Due to the non-linearity of electric vehicle\u2019s some parts (such as motors),tires will occur serious non-linearity during the driving process of electric vehicles.The model of 4WIMD electric vehicle is as shown in Fig.1,vehicle dynamics model is the seven-degree-of-freedom(DOF),and Vx,Vy and \u03b3 are the longitudinal speed,lateral speed and yaw rate.Fxi, Fyi, Fzi(i=fl,rl,fr,rr)are the longitudinal,lateral and vertical forces.Mz is yaw torque.Iz is centroid moment of inertia. lf is distance from O to front axle.lr is distance from O to rear axle.l = lf + lr is distance between 2 axles.\u03b4 is steering angle of vehicle front wheels.\u03b2 is vehicle body sideslip angle.m is vehicle quality.tr is front(rear) axle track,which is built on the following assumptions:(1)The front wheel and rear wheel have the same track pitch,and the suspension characteristics are ignored;(2)Longitudinal,lateral pitch and roll angles of the vehicle are ignored",
            " The simplified formula of the lateral force is the function of vertical load and the tire side declination,and also expresses the nonlinear characteristics of tire lateral force.Fig.3 shows the contrast curve of the side slip force between the magic formula and the simplified model when the parameter C changes dynamically.Road adhesion coefficient is good,and the fitting parameters in the magic formula are derived from literature[21].D\u2019s value is 0.15,and E\u2019s value is 1.3.It can be seen that with the constant change of parameter C,the simplified formula curve can approach the magic formula. The vehicle model is shown in Fig.1.It includes longitudinal,lateral and yaw motions about the vertical axis,and the front wheels are drive wheels.According to the above formula,the state vectors x consisting of vehicle sideslip angle,yaw angular velocity and front and rear tire lateral forces are defined as follows: x = [\u03b2, \u03b3, Fyf , Fyr] T = [x1, x2, x3, x4] T (8) The observation vector y consisting of the yaw rate and the lateral forces of the front and rear wheels are as follows: y = [\u03b3, Fyf , Fyr] T = [y1, y2, y3] T (9) The steering angle and the longitudinal driving forces of the four wheels constitute the input vector u as follows: u = [\u03b4, Fxfl, Fxfr, Fxrl, Fxrr] T = [u1, u2, u3, u4, u5] T (10) The inputs of the system are the front wheel steering angle,the left front tire force,the left rear tire force,the right front tire force,the right rear tire force,and the state is estimated by the state equation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure45.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure45.3-1.png",
        "caption": "Fig. 45.3 Various components of the developed ginger plant",
        "texts": [
            " The design of a power-operated mini ginger planter was conceptualized and fabricated. The developedmachinewas attachedwith a power tillerwhich is awalk-behind type. The design constitutes hopper (1), metering bucket (2), chain (3), sprocket (4), metering device covering unit (5) frame (6), furrow opener (7), furrow covering device (8), and furrow wheel (9). The developed planter consists of two furrow wheel and a horizontal frame that supports all other components. The various components of the ginger planter were shown in Fig. 45.3. The ginger planter was tested in the laboratory and the field conditions. In the laboratory testing, evaluated parameters like a seed to seed spacing and depth of furrow opened. However, in the field testing evaluated various parameters like mean seed missing (%), mean bucket filling (%), percentage of physical damage (%), theoretical field capacity (ha/h), actual field capacity (ha/h), field efficiency (%), depth of plantation (cm), and seed spacing (cm). The developed ginger planterwas tested in the laboratory for performance evaluation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002161_cdc42340.2020.9303889-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002161_cdc42340.2020.9303889-Figure1-1.png",
        "caption": "Fig. 1. Dynamics model: a point mass my is connected to a quadrotor via a rigid link of the mass ml and length l. Its configuration is defined by the location of the point mass y \u2208 R3, the direction q \u2208 S2 of the link from the quadrotor towards the point mass, and the attitude of the quadrotor R \u2208 SO(3).",
        "texts": [
            " Finally, the proposed control system is validated through a numerical simulation and an indoor flight experiment where the payload follows a nontrivial trajectory in the three-dimensional space. In short, the main contributions of this paper are threefold: accounting the effects of the mass distribution of the link in the dynamic model of aerial transportation; proposing a new geometric tracking control system that incorporates the mass of the link while compensating disturbance; and validating the geometric tracking control system with flight experiment. Consider a quadrotor UAV that is connected to a point mass my via a rigid link as shown at Figure 1. Define an inertial reference frame {~e1, ~e2, ~e3} and a body-fixed frame {~b1,~b2,~b3}. The body-fixed frame is attached to the quadrotor and its origin coincides with the mass center of the 978-1-7281-7447-1/20/$31.00 \u00a92020 IEEE 201 Authorized licensed use limited to: Carleton University. Downloaded on June 03,2021 at 02:20:20 UTC from IEEE Xplore. Restrictions apply. quadrotor, where the first axis point towards a rotor, and the third axis points downward. The location of the quadrotor is denoted by x \u2208 R3 in the inertial frame, and its attitude is R \u2208 SO(3) = {R \u2208 R3\u00d73 |RTR = I3\u00d73, det[R] = 1}"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000547_iros40897.2019.8967922-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000547_iros40897.2019.8967922-Figure7-1.png",
        "caption": "Fig. 7: Simulation results of the proposed planner for the aerial pick-and-place for five different environment settings. The red line indicates the trajectory of the body position of the multi-rotor. Green lines show the trajectories of tip of the link. The number of total nodes in simulations (a),(b), and (c) is 10,000, and it is 100,000 in simulations (d) and (e).",
        "texts": [
            " The target object is cylinder-shaped and located on the box. The resulting trajectories from the proposed planner are also shown in Figs. 7a\u20137e. The red lines indicate the trajectories of the body position for multi-rotor. Green lines show the trajectories of the tip of the link. In TABLE I, the resulted costs and computational times for each case are listed. The numbers of sampling nodes are selected considering the size of the volume of environments. In TABLE II, we listed the detailed computational times in each process for the fifth case (Fig. 7e). The comparison of cost and computation time between the best waypoint and the worst waypoint is also shown in TABLE II. With the proposed merging process, we are able to find the best pose of the aerial manipulator grasping the object. For a pick-and-place task, we validate the reliability of the generated path with autonomous flight experiments. In this section, we describe the experimental set-ups and results. Our aerial manipulator is composed of a Flame Wheel F550 from DJI and a 2-DoF robotic arm which consists of MX series from ROBOTIS"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000958_psc49016.2019.9081498-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000958_psc49016.2019.9081498-Figure4-1.png",
        "caption": "Fig. 4. The magnetic flux density distribution of FSPM motor",
        "texts": [
            " Then, ReliefF averages their contribution for updating the quality estimation W[F] for all features F, that include wavelet and statistical features. The ReliefF algorithm is as Fig. 3 [27]. IV. SIMULATION VALIDATION In order to validate the proposed approach, the FEM of the FSPM is simulated for different conditions. The studied cases of the PM of the FSPM, are healthy state, 20% demagnetization of one PM, 40% demagnetization of one PM, and 80% demagnetization of one PM. The flux density distribution of the FSPM motor under healthy condition is shown in Fig. 4. For one load condition and considering one of the PMs, the simulation results of electromagnetic torque is shown in Fig. 5 for the mentioned case studies. The FFT of the obtained electromagnetic torque from FEM is derived and then are applied to ReliefF. The harmonic components for each state construct a 20 component set which comprises harmonic components of 700 Hz, 1400 Hz, 2100 Hz, \u2026 , 14000 Hz. These components are ranked by ReliefF and result 96 Authorized licensed use limited to: University of Exeter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002726_j.ymssp.2021.107837-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002726_j.ymssp.2021.107837-Figure1-1.png",
        "caption": "Fig. 1. The 3D illustration of an SEA.",
        "texts": [
            " In addition, rigorous stability analysis is also conducted on the derived closedloop control system. The effectiveness of the proposed controller is verified in the experiments in Section 4. At last, the conclusions are given in Section 5. Before deriving SEA\u2019s dynamical model and its controller, we introduce some notations here. Define a set Ni:j \u00bc fi; i\u00fe 1; . . . ; jg, for integers i and j satisfying 0 6 i 6 j. For any integer i P 0, let \u00f0i\u00de denote the ith-order time derivative of any variable or function . As depicted in Fig. 1, the dynamical model of the SEA can be illustrated by analogy of a two-mass-spring-damper system whose mathematical model can be expressed as Jl\u20achl \u00fe Bl _hl \u00femlgL sin\u00f0hl\u00de \u00bc Ks\u00f0hm hl\u00de \u00fe nl\u00f0t\u00de Jm\u20achm \u00fe Bm _hm \u00fe Ks\u00f0hm hl\u00de \u00bc u\u00fe nm\u00f0t\u00de \u00f01\u00de where hl; _hl and \u20achl are angular position, angular velocity and angular acceleration of the load side of the SEA, respectively; Jl and Bl denote the inertia and viscous friction coefficient of the load side, respectively; ml denotes the total mass of the load side; g is the acceleration of gravity; L is the distance from the axis of rotation to the center of mass of the load; Ks is the stiffness coefficient of the torsional spring in the SEA; nl\u00f0t\u00de represents other unknown disturbance acting on the load side that is independent of the system states; hm; _hm; \u20achm; Jm;Bm and nm\u00f0t\u00de are corresponding physical quantities for the motor side; u denotes the control input torque provided by the geared motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000989_s40430-020-02392-5-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000989_s40430-020-02392-5-Figure2-1.png",
        "caption": "Fig. 2 Planetary gear transmission",
        "texts": [
            " These machines mainly use planetary gear box because they provide huge torque which is used for the bending of pipes. CNC pipe bending machines as shown in Fig.\u00a01 can bend metal pipes from sections just a few millimeters to several meters long. Pipe bending machines use various types of dies to produce the specified shape of pipe during the process of bending. Sometimes it may use heat treatment equipment to dissipate the heat from the component as it affects the surface and lifetime of the component. A planetary gear box (also called epicycle gear train) as shown in Fig.\u00a02 mainly contains three gears named as sun gear, planet gear and ring gear. They are situated in such a manner that the center of planet gear revolves around the center of the sun gear. Arm is the connecting rod which holds sun and planet gear. During the process of meshing of sun and planet gear, their pitch circles rolls without slipping. On the pitch circle of planet gear, a point traces a curve named as epicycloid curve. In general case, the number of planet gear is 3, but it may vary depending on requirement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001535_j.jterra.2020.08.003-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001535_j.jterra.2020.08.003-Figure3-1.png",
        "caption": "Fig. 3. Tomcar TM27 side-by-side ATV model.",
        "texts": [
            " The model of the wheeled vehicle consisted of 5 bodies: a chassis and 4 wheels. The front suspension was double wishbone, built of 4 stiff springs and dampers, and the spring and shock absorber of the suspension. The rear suspension was a trailing arm, built of 2 stiff springs and dampers, and the spring and shock absorber of the suspension. Powertrain and cruise control operated external forces on the model. The steering control operated a constraint on the model (therefore, the steering torque could not be calculated). Fig. 3 shows some of these parts on the Tomcar TM27 side-by-side ATV wheeled vehicle. In this section, modeling of some parts of the vehicle will be described. The chassis is a body with mass of 612.9 kg and inertia of: Ixx = 187.5 kgm2, Iyy = 523.5 kgm2, and Izz = 636 kgm2 at the center of mass. A reference coordinate system is located above the center of the rear suspension at 0.381 m above the ground. The reference coordinate\u2019s orientation is same as the axis system of CG (Fig. 3). The gravity center is located at 0.784 m front to the reference coordinate system and 0.094 m above the reference coordinate system. Table 2 shows the connection points of the stiff springs and dampers to the body relative to the reference coordinate system. Table 2 Location of the connection points of the suspensions in the reference coordinate system. Suspension Connection point X [m] Y [m] Z [m] Front suspension Lower arm \u2013 front connection 2.1755 0.223 0.103 Lower arm \u2013 rear connection 1.8335 0",
            " The results presented here describe the travel of the vehicle at a velocity of 5 m/s. In all the illustrations, DBD is compared with simulation of multibody dynamics using the Siemens\u2019 VL software. The kinematics relative to the horizontal position of the vehicle are described in Fig. 10. The illustration on the right describes the position of the center of gravity of the chassis on the Z axis, and the illustration on the left describes the pitch angle (negative value describe pitch up according to the reference system as shown in Fig. 3). The figure shows that there was a good match between the two computation methods regarding the vehicle\u2019s kinetics, with a maximum error of less than 4% between the different simulations. Fig. 11 presents a comparison between the forces vs. time calculated using the two methods. The figure on the right depicts the forces of the rear spring; the figure on the left depicts the emerging power in one of the bushing front suspensions that connected the vehicle chassis. As seen, a good match was obtained regarding the former, with a maximum error of less than 3% between the simulations using different methods"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002755_tmag.2021.3074935-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002755_tmag.2021.3074935-Figure8-1.png",
        "caption": "Fig. 8. Rated load radial force density distributions of Models 1\u20134. (a) Model 1: 8/6/2. (b) Model 2: 5/3/2. (c) Model 3: 7/3/2. (d) Model 4:7/6/1.",
        "texts": [
            " 7 shows the flux distributions on rated load (Idc = 13.44 A and Iac = 13.44 A) of models. Similar to no load distribution, the flux distribution of Model 1 is centrosymmetric and that of model 4 is symmetric while for Models 2 and 3, it is asymmetric. Local saturation occurs in the stator tooth in Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 24,2021 at 17:02:44 UTC from IEEE Xplore. Restrictions apply. Models 2 and 3 and the stator yoke in Mode 4 due to asymmetric flux distribution. Fig. 8 exhibits the radial force density of four models at the same rotor position as in Fig. 7. As can be seen, unbalanced magnetic force exists in Models 2\u20134, which might likely produce vibrations and noise. With the rotating exciting fields and rotating armature MMF, the electromagnetic torque ignoring the iron saturation effect can be obtained with the following equation: Tem = 2\u03c0rgleff B A \u221d 2\u03c0rgleff Idc Iac N2 a (8) where A is the armature electrical loading. It can be seen that the torque is proportional to the square of the number of turns, and the biased dc component and the ac component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002176_0142331220981430-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002176_0142331220981430-Figure2-1.png",
        "caption": "Figure 2. Geometry for an attacker and its neighbors.",
        "texts": [
            " Ri is the relative distance between the target and the i-th attacker, and Vi is a positive constant that describes the initial speed of the ith attacker. The terms gi and gT are the heading angles of the attacker and the target. li is the i-th attacker\u2019s LOS angle in the inertial reference frame. ji is the bearing angle between the direction of the i-th attacker\u2019s velocity and the i-th LOS, and the term fi is the bearing angle between the direction of the target\u2019s velocity and the i-th LOS. Multi-attackers with uniform velocity are denoted as nodes V= 1, ::,N . For simplicity, a geometry is described for the i-th attacker and its neighbors in Figure 2. uii+ 1 is the angle from the i-th LOS to the i+1-th LOS, and rii+ 1 is the relative distance between the i+1-th attacker and the i-th attacker. VT is a positive constant that describes the initial speed of the target. The connection between the angles gT , fi and l\u0302i is formed in Figure 3. It is noteworthy that the attacker and the target have their accelerations perpendicular to the directions of their speeds, respectively, which means their speeds are pre-determined and the directions of the speeds change constantly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000189_ecce.2019.8912698-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000189_ecce.2019.8912698-Figure5-1.png",
        "caption": "Fig. 5. Wall shear stress distribution on the slotted rotor at \u2126=20000 rpm",
        "texts": [
            " Wall Shear Stress The shear stress is a resistance to fluid movement caused by friction which is related to the fluid viscosity. A fluid at rest cannot resist shearing forces and deforms continuously, however small they are. The resistance to the action of shearing forces in a fluid appears only when the fluid is in motion. The dimensionless skin friction coefficient Cf should be determined to estimate the windage losses. As defined in Eq. 2 this coefficient is defined by the ratio of the shear stress in the rotor surface to the dynamic pressure pd. As shown in Fig. 5 the wall shear stress is more important on the edge between surfaces S4 & S5 where the airflow attacks the rotor. In this region, the air dynamic contributes to increase the shear stress on the rotor surfaces. C. Influence of the rotation speed on skin friction coefficient As shown in Eq. 2, the skin friction coefficient Cf is expressed as a function of the wall shear stress i which is calculated by the CFD model. The coefficient can be calculated for each subsurface of the rotor from S1 to S5. As shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001075_iccar49639.2020.9108090-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001075_iccar49639.2020.9108090-Figure2-1.png",
        "caption": "Figure 2. Rigid five-bar model. The reference coordinate system (coordinate system 0) is defined at the heel of standing leg.",
        "texts": [
            "2 is used to process experimental data and simulate the human walking, then the angle and torque data of the hip, knee and ankle joints in the single gait cycle in the sagittal plane can be obtained. human body. The human body is considered as a plane multibody system, and only planar motion in the sagittal plane is studied. The left and right legs are simplified as rods connected by hinges, and the mass of the torso includes the upper limbs and head. Kinematic model can be developed according to the Denavit-Hartenberg (D-H) representation method of robotic kinematics. As shown in Fig. 2, the reference coordinate system, local coordinate system and relevant parameters of the five-bar model are defined. The D-H parameters of the model are listed in Table 1. When the subject walks, the swinging foot is regarded as the end-effector, and the relationship between the coordinate of the swinging foot in the reference coordinate system and the joint angles can be given by kinematic equations. The local coordinate system i (i=1, 2, 3, 4, 5, 6) is obtained by translation and rotation of the reference coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001530_0954406220957369-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001530_0954406220957369-Figure1-1.png",
        "caption": "Figure 1. 3D sectional view and unwrapped view of bearing with (a) Rectangular pocket, (b) Trapezoidal pocket, (c) Elliptical pocket, (d) Parabolic pocket, and (e) Schematic of journal bearing with the coordinate system.",
        "texts": [
            " It is observed that pocketed bearing may increase the threshold speed of instability of rotor in comparison with smooth bearing if proper extent of pocket is chosen. Among all pocket shapes, the rectangular shape of the pocket gives the best stability characteristics. Therefore, it may be concluded that bearings may be given a rectangular pocket to increase the stability of rotors. A three-dimensional view of pocketed journal bearing, unwrapped view of bearing, and schematic with coordinate system is shown in Figure 1. Haff14 developed simplified theory for modelling the grain flow from a continuum perspective. Haff\u2019s theory assumes grain particles as identical spheres and the average separation (s) between the particles to be far lesser than the grain diameter (d) so that the average density (q) may be assumed constant. Furthermore, cohesion of particles and spin are also neglected. Apart from bulk flow velocity (V) each particle has an average fluctuation velocity ( v) also known as granular temperature. Utilising the conservation laws of mass, momentum, and energy, governing equations were derived",
            "3,26 Thus, integrating the momentum equations (5) and (7) and applying the boundary conditions, (i) at y\u00bc 0, ux\u00bc 0, uz\u00bc 0, and (ii) y\u00bc h, ux\u00bcU, uz\u00bc 0, the relation for flow velocities are obtained as, ux \u00bc 1 2g @p @x y2 hy \u00feU h y (13) uz \u00bc 1 2g @p @z y2 hy (14) Now, the volume flow rate per unit width in x and z directions can be obtained by integrating the flow velocity as follows Qx \u00bc Z h 0 uxdy \u00bc h3 12g @p @x \u00feUh 2 (15) Qz \u00bc Z h 0 uzdy \u00bc h3 12g @p @z (16) Substituting equations (13) and (14) in the continuity equation (2) and integrating along film thickness, the Reynold\u2019s equation is obtained as; @ @x 1 6 t a wh2 @\u00f0lnp\u00de @x \u00fe @ @z 1 6 t a wh2 @\u00f0lnp\u00de @z \u00bc U 2 @h @x \u00fe @h @t (17) where w \u00bc ffiffiffiffiffiffiffiffi k=w p B ekh \u00fe e kh 2\u00f0 \u00de= ekh e kh\u00f0 \u00de Film thickness variation in pocketed journal bearing is expressed as follows: h \u00bc 1\u00fe e cos h /\u00f0 \u00de\u00bd \u00fe D h (18) where h \u00bc h=Cr, and D h is the film thickness component due to the presence of a pocket. The relations for D h applicable to different shape of pockets (refer Figure 1) are written as below: (i) Rectangular pocket D h \u00bc dp ; if x a; 1 2 rz z 1 2 \u00fe rz 0 ; if x > a; z < 1 2 rz ; z > 1 2 \u00fe rz 8>>< >>: (19a) (ii) Trapezoidal pocket D h \u00bc dp; if x a; 1 2 rz \u00fe b xtan d\u00f0 \u00de z 1 2 \u00fe rz b xtan d\u00f0 \u00de 8>>< >>: 9>>= >>; 0 ; if x > a; z < 1 2 rz \u00fe b xtan d\u00f0 \u00de ; z > 1 2 \u00fe rz b xtan d\u00f0 \u00de 8>>< >>: 9>>= >>; 8>>>>>>>< >>>>>>>: (19b) (iii) Elliptical pocket D h \u00bc dp; if x a 2 \u00fe z 0:5 rz 2 \" # 1 0; if x a 2 \u00fe z 0:5 rz 2 \" # > 1 8>>>< >>>: (19c) (iv) Parabolic pocket D h \u00bc dp; if x a \u00fe z 0:5 rz 2 1 \" # 0 0 ; if x a \u00fe z 0:5 rz 2 1 \" # > 0 8>>>< >>>: (19d) where b \u00bc pD=L, and rz \u00bc rz=L Film force acting on the journal is resolved in vertical (X) and horizontal (Y) directions (refer Figure 1(e)). The expressions for force components in \u2018X\u2019 and \u2018Y\u2019 directions are written below: FX \u00bc Z L 0 Z 2p 0 pRcosh dhdz (20a) FY \u00bc Z L 0 Z 2p 0 pRsinhdhdz (20b) Resultant internal force (Fint) on the journal due to film pressure is evaluated using the following relation: Fint \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 X \u00fe F2 Y q (21) Shear force (Fs) is computed using the following relation:3 Fs \u00bc Z L 0 Z 2p 0 h 2R @p @h \u00fe g U h Rdhdz (22) Coefficient of friction (f) is calculated using follow- ing relation: f \u00bc Fs Fint (23) Dynamic characteristics of rotor-bearing system To study the dynamic characteristics of a rotorbearing system, it is important to know the stiffness and damping coefficients of the bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003415_9781119352181.ch4-Figure4.7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003415_9781119352181.ch4-Figure4.7-1.png",
        "caption": "Figure 4.7 Circular slot.",
        "texts": [
            "25) Hence the fraction of the total turns enclosed by the flux line is n (x) ns = A (x) Ac = [x(y + bs)]\u22152 [d3(bs + b2)]\u22152 = y + bs b2 + bs ( x d3 ) (4.26) The specific permeance of the conductor-occupied region is, from equation (4.11), p3 = \ud835\udf070 d3 \u222b 0 ( y + bs b2 + bs )2( x d3 )2 dx y (4.27) which yields ultimately p3 = \ud835\udf070 d3 bs \u23a1\u23a2\u23a2\u23a2\u23a3 \ud835\udefd2 \u2212 \ud835\udefd4 4 \u2212 ln \ud835\udefd \u2212 3 4 (1 \u2212 \ud835\udefd)(1 \u2212 \ud835\udefd2)2 \u23a4\u23a5\u23a5\u23a5\u23a6 (4.28) where \ud835\udefd = b2 bs The total permeance of the coffin-shaped slot is simply the sum of p012 and p3 defined respectively by equations (4.22) and (4.28). As a final example consider the circular slot of Figure 4.7. Again this is a widely used configuration for rotor slots particularly in larger machines where copper rods are inserted in the slots and soldered together at each end of the rotor to form a squirrel cage. The specific permeance of the slot opening is, almost by inspection p0 = \ud835\udf070d0 bo Over the circular portion, the length of an infinitesimal tube of flux at a distance x from the bottom of the slot can be expressed as y = 2r sin \ud835\udefc (4.29) where \ud835\udefc = cos\u22121 ( r \u2212 x r ) (4.30) or, x = r \u2212 r cos \ud835\udefc so that dx = r sin \ud835\udefcd\ud835\udefc (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002504_j.ijmecsci.2021.106392-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002504_j.ijmecsci.2021.106392-Figure8-1.png",
        "caption": "Fig. 8. Mechanical model of the torus.",
        "texts": [
            " Therefore, the set of generalized = ( \ud835\udc65 \ud835\udc36 \ud835\udc66 \ud835\udc36 \ud835\udf13 \ud835\udf19 \ud835\udf03 )T . he absolute (inertial) frame corresponds to body 1, whereas the torus y \ud835\udc4e and \ud835\udc4f, respectively, and the ratio = \ud835\udc4e \ud835\udc4f ereinafter will be referred to as the aspect ratio. In order to make furt elation has been considered: + \ud835\udc4f = \ud835\udc45. I e of the torus, and contains the hoop of equivalent radius. The contact point o ntre of the torus tube, and \ud835\udc3c is the lowest point of the equivalent hoop. \u2032\u27e9, \u27e8\ud835\udc4b \u2032\u2032\ud835\udc4c \u2032\u2032\ud835\udc4d \u2032\u2032\u27e9 and the body frame 2 \u27e8\ud835\udc4b 2 \ud835\udc4c 2 \ud835\udc4d 2 \u27e9 shown in Fig. 8 is given by E an be computed as: \ud835\udc93 (91) w \ud835\udc93 (92) without slipping. Thus, the velocity of the contact point can be computed f \ud835\udc97 (93) N nholonomic constraints are: \ud835\udc6a (94) G e system is \ud835\udc5b \ud835\udc54 = 3 . The equations of motion of the torus constitute a nonlinear i the nonholonomic constraints, results in the following index-1 DAE system: ( (95) w ) ) sin 2 ( \ud835\udf19) \ud835\udc6b \ud835\udc78 os ( \ud835\udf13 ) cos ( \ud835\udf19) \u0307 sin ( \ud835\udf13 ) cos ( \ud835\udf19) ?\u0307? sin ( 2 \ud835\udf19) n ( \ud835\udf19) cos ( \ud835\udf19) \u239e \u239f \u239f \u239f \u239f \u239f \u23a0 , \ud835\udc85 (96) I r and a vertical axis perpendicular to the torus centre, respectively: \ud835\udc3c (97) n this way, the plane \ud835\udf0b\ud835\udc5a shown in Fig. 8 corresponds to the middle plan f the torus with the ground is denoted by \ud835\udc36; point \ud835\udc43 represents the ce As in the hoop, the orientation of the intermediate frames \u27e8\ud835\udc4b \u2032\ud835\udc4c \u2032\ud835\udc4d qs. (70) . The absolute position vectors of \ud835\udc36 and the centre of mass \ud835\udc3a c \ud835\udc36 = ( \ud835\udc65 \ud835\udc36 \ud835\udc66 \ud835\udc36 0 )T , \ud835\udc93 \ud835\udc3a = \ud835\udc93 \ud835\udc36 + \ud835\udc93 \ud835\udc36\ud835\udc43 + \ud835\udc93 \ud835\udc43\ud835\udc3a , ith \ud835\udc36\ud835\udc43 = ( 0 0 \ud835\udc4e )T , \ud835\udc93 \ud835\udc43\ud835\udc3a = \ud835\udc79 \u2032\u2032(0 0 \ud835\udc4f )T . The nonholonomic constraints arise from the assumption of rolling rom Eq. (72) , obtaining: \ud835\udc36 = \u239b \u239c \u239c \u239c \u239d \ud835\udc63 \ud835\udc36 \ud835\udc65 \ud835\udc63 \ud835\udc36 \ud835\udc66 \ud835\udc63 \ud835\udc36 \ud835\udc67 \u239e \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239d ?\u0307? \ud835\udc36 \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) \u0307\ud835\udf03 cos ( \ud835\udf13 ) \u2212 \ud835\udc4e "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001662_012010-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001662_012010-Figure4-1.png",
        "caption": "Figure 4. Deflection of the conductor presented by two methods.",
        "texts": [
            " The approximate solution (12) will also satisfy all the boundary conditions, but in general, will not satisfy (3) identically since there remains an error defined by 2( , ) ,T Te x t c \u2032\u2032= \u2212U U \u03a6 \u03a6 (13) which one can force it to have a zero projection on the chosen functions ( )x\u03a6 . 0 ( , ) d 0. l ne x t x =\u222b \u03d5 (14) Substituting the expression of the error from equation (13) in equation (14), and rewriting, we have: 0 2 0 , d , d , . l T l T x c x p \u03a6\u03a6 \u03a6\u03a6 \u03a6 + = = \u2032\u2032= = \u222b \u222b MU KU f M K f  (15) The equation (15) can be transited to a discrete model of a wire with a moving load to an ODE system (linear, with constant coefficients). System (15) is represented in the form of first-order equations: 1 . ( )\u2212    U = \u0396 \u0396 =M f -KU   (16) The result is shown in figure 4. Parameters are the same as the previous section. ISPCIET 2020 IOP Conf. Series: Materials Science and Engineering 939 (2020) 012010 IOP Publishing doi:10.1088/1757-899X/939/1/012010 In this section, we consider the resistance force from the wind on the electrical line. To take into account the distributed forces of external resistance in the string equation (3), the new form will be obtained [17\u201320]: ( , )u Tu bu q x t\u2032\u2032\u2212 + = \u03c1 (17) To solve the equation (17) using the Galerkin method, the following change can be made in the equation (15): 0 2 0 , d , d , "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001812_sces50439.2020.9236725-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001812_sces50439.2020.9236725-Figure1-1.png",
        "caption": "Figure 1. CAD model of the 3-PSS Parallel Manipulator",
        "texts": [
            " The results and discussion are portrayed in section 4. The conclusion and future scope of work are given in section 5, respectively. II. THE ARCHITECTURE OF THE 3-PSS Parallel Manipulators are closed-loop mechanisms consisting of a moving platform connected by more than one leg. Due to the multiple connections to the moving platform, the precision and accuracy achieved will be higher compared to that of serial manipulators [10]. The CAD model of the 3- PSS parallel manipulator considered in this study is shown in Fig. 1. The manipulator consists of three limbs, with each limb composed of one active prismatic joint and two passive spherical joints. The three active sliders are actuated along the horizontal direction with each corner of the equilateral triangle-shaped mobile platform connected to each slider via a rigid link having passive spherical joints at each end. The global frame {O} is fixed at the top centre, as shown in Fig.1, and the moving frame is located at the centre of the platform. Let Li be the slider distance of the ith leg measured along the x-axis of the global frame; c be the perpendicular distance between two consecutive sliders and l be the length of the rigid link. The 3- DoF of the 3-PSS manipulator corresponds to the two linear motions and one angular rotation of the end effector platform. These three degrees are controlled by manipulating the three active sliders. Therefore, this design does not have any redundant DoF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000569_j.nahs.2020.100876-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000569_j.nahs.2020.100876-Figure1-1.png",
        "caption": "Fig. 1. Underlying IPN of the LHS. The firing sequences t1t3t5t7t9 and t2t4t6t7t9 produce the same output changes, i.e., both sequences are indistinguishable, thus the IPN is not observable.",
        "texts": [
            " Theorem 2 also includes the case of LHS in which the discrete state can be determined by only using discrete observations [29]. Moreover, the results here presented includes a more general case when the discrete state is not observable by using only discrete information and the continuous state is not observable by using only continuous information, for the autonomous case [33], with known inputs and with (unbounded) unknown inputs; providing more relaxed sufficient conditions. 4.6. Example Consider a LHS that consists of the IPN of Fig. 1 and the collection of the LS\u2019s described in Table 1 (each place pi of the IPN is associated to the continuous system \u03a3i = (Ai, Bi, Fi, Ci)). The IPN is an ordinary safe and live state machine, having label types {A, B, C,D} with output matrix \u03a6 = \u23a1\u23a2\u23a30 1 1 0 0 0 0 0 0 0 0 1 1 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 \u23a4\u23a5\u23a6 . This system could be representing the operation of a machine, where \u03a31, \u03a36 describe the setting up and shutting down dynamics of the machine, respectively. \u03a33, \u03a35 describe two normal dynamical operation modes of the machine, while \u03a32, \u03a34 describe two faulty dynamical operation modes",
            " The normal operation sequence is \u03a31\u03a33\u03a35\u03a38, while the faulty operation sequence is \u03a31\u03a32\u03a34\u03a37. Thus, the observer can be used to detect if a fault occurred during the machine operation. Let us firstly investigate the observability of the IPN . The following sequences describe minimal T-semiflows: \u03c31 = t1t3t5t7t9, \u03c32 = t2t4t6t7t9, \u03c33 = t1t3t5t8t10 and \u03c34 = t2t4t6t8t10. Sequences \u03c31 and \u03c32 produce the same output changes. Similarly, sequences \u03c33 and \u03c34 produce the same output changes. Thus, due to the existence of these indistinguishable sequences, the IPN of Fig. 1 is unobservable. Furthermore, notice that, even if the marking of the net were known to be at p1, the knowledge of the marking would be lost after the firing of any transition in the conflict {t1, t2}, since their firings produce the same output change. On the other hand, almost all the LS\u2019s are indistinguishable between them for the case of unknown initial condition, excepting \u03a35 and \u03a36 in which Fi = 0. Consequently, it is not possible to determine the active LS by only using the continuous output information"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000359_icaaid.2019.8935001-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000359_icaaid.2019.8935001-Figure1-1.png",
        "caption": "Fig. 1. The view of the aerial vehicle [1]",
        "texts": [
            " The mathematical model of the system is addressed in the next section. In the third section, two recent control strategies, namely, model-free intelligent PID (iPID) control and type-2 fuzzy adaptive PID control is discussed. The control methods are investigated according to several quantitative measures. Furthermore, the pros and cons of each control strategy are addressed. Finally, the results and contributions are given. The structure of quadrotor and the attached coordinates are depicted in Fig.1. The dynamic modelling is presented in several studies, for instance, interested readers are referred to some recent results of system modelling and identification studies such as [26], [27], [28]. The dynamical equations are shown in (1) [29]. mx\u0308 =(c\u03c6 s\u03b8 c\u03c8 + s\u03c6 s\u03c8)U1 my\u0308 =(c\u03c6 s\u03b8 s\u03c8 \u2212 s\u03c6 c\u03c8)U1 mz\u0308 =\u2212g+(c\u03c6 c\u03b8)U1 \u03c6\u0308 =\u03b8\u0307 \u03c8\u0307 (Iyy \u2212 Izz) Ixx + Jr Ixx \u03b8\u0307\u03a9d + l Ixx U2 \u03b8\u0308 =\u03c6\u0307 \u03c8\u0307 (Izz\u2212Ixx) Iyy \u2212 Jr Iyy \u03c6\u0307\u03a9d+ l Iyy U3 \u03c8\u0308 =\u03b8\u0307 \u03c6\u0307 (Ixx\u2212Iyy) Izz + l Izz U4 (1) The rotational speeds of the four motors and the thrusts are represented by \u03c9i and Ti, respectively "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure15-1.png",
        "caption": "Fig. 15. (a) Rotor center at a simulated position and stator body at the nominal position; all the components being subjected to contact forces. (b) Indentation distribution on the roller.",
        "texts": [
            " The flow through this gap is considered as the case of flow through a circular pipe of equivalent gap length (Fig. 14). This equivalent gap length (lt) was evaluated as distance between points on the gears that satisfy h t \u00bc ht 1\u00fe \u00f0 \u00de where was found to be 0:36 based on both experiments and simulations individually [25]. It should be noted that the stator presented in Fig. 13(a) is a combination of stator body and rollers. To evaluate the contact forces more accurately, the contact between the rollers-stator body is also modelled as shown in Fig. 15(a). For a given roller, the net indentation between the stator and the rotor is evaluated (Fig. 13(b)), and this indentation, d, is distributed between the rotor-roller contact, d1, and the roller-stator body contact, d2, as shown in Fig. 15b (inset of a roller in Fig. 15 (a)) such that d \u00bc d1 \u00fe d2 and the forces (highlighted by arrows) evaluated at both the contacts are same so that the net force on the roller is nullified. A one-dimensional root finding algorithm [26] is employed to find the value of d1 and consequently, d2. Apart from the magnitude of the contact forces, both the direction and the arm-length of these forces with respect to the rotor center are also evaluated. Upon evaluation of these quantities at all the possible contact points, the values are stored in a lookup table for subsequent use"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000496_oceans40490.2019.8962582-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000496_oceans40490.2019.8962582-Figure1-1.png",
        "caption": "Fig. 1. Reconfigurable voodoo-waveguide structure. (a) Design plan view. The shaded areas represent the position of the holes (nominal dimensions are",
        "texts": [
            " This is of particular relevance when this technique is used by end-users for routine verification of waveguide VNAs. The proof-of-concept verification process that is described clearly demonstrates that this new approach to calibration verification can quickly reveal whether a calibrated VNA is operating within expected performance metrics. The behavior of a reconfigurable \u201cvoodoo-waveguide\u201d is first investigated theoretically using a full-wave simulator (CST MWS). The design of the reconfigurable voodoo-waveguide structure chosen on this occasion is shown in Fig. 1, having five pins/holes to demonstrate the single-component verification concept. For example, with reference to Fig. 1(a), hole is placed at the center of the waveguide (along both the - and -directions), mm, mm, mm, mm, and mm). (b) Machined structure with no pins inserted and (c) machined structure with two pins inserted all the way through. where the electric field is at a maximum for the fundamental mode, resulting in the highest level of interaction when a pin is inserted. Similarly, and are located in-line with (along the -direction), with an offset and from the adjacent sidewalls, respectively. As a result, different levels of interference can be achieved to create predetermined changes in the reflection and transmission coefficients",
            " Polar plot responses exhibit a spiral frequency response (for both reflection and transmission measurements), which is unique to the reconfigurable voodoo structure. This demonstrates that this single-component verification kit should be sufficient for most practical applications. Having values spread out across most of the scattering ( )-parameter planes provides a comprehensive verification for the VNA\u2019s reflection and transmission measurements, as will be discussed in detail in the following sections. Next, a prototype device based on the initial design is fabricated, as shown in Fig. 1(b) and (c). The holes are made by electrical discharge machining using a computer-controlled process. A spark from a copper rod placed in close proximity to the aluminum waveguide surface repeatedly vaporizes a small part of the waveguide near the copper rod. The vertical line of holes , , and is nominally midway between the waveguide end ports. Fig. 1(b) and (c) shows the finished rectangular waveguide, having internal aperture dimensions mm (corresponding to WR-15 [14]), with symmetrical holes drilled on both broad walls. Through the holes, high-speed steel cylindrical pins can penetrate though the waveguide walls. A. Reconfigurable Structure Dimensional Measurements In order to verify successful manufacture of the reconfigurable voodoo-waveguide structure, it is important to check the positions of the pin insertion holes at the micrometer level of accuracy",
            " The position of the waveguide\u2019s inner sidewalls, with respect to the outer sidewalls, was determined using a Zeiss Universal Precision Measuring Center (UPMC) CMM fitted with a \u201cT\u201d-shaped stylus array [see Fig. 3(b)], each stylus having a 600-\u00b5m-diameter ball-ended tip. In this case, a second local coordinate system (CS2) was set up using designated waveguide datum features. This was then used as the global coordinate system. The Port 1 aperture was defined as the plane, and one internal sidewall defined as the plane, as indicated in Fig. 1(a), with the origin at . The latter is designated as the wall adjacent to holes and . Since it was not possible to make contact with the inside of any sidewall, along the length of the waveguide, 24 points corresponding to the inside of this sidewall were determined at each end of the waveguide with Gaussian best fit planes fitted to the data. A mean plane was constructed from the part planes at both ends to define the plane. The external position of the sidewall, with respect to the plane, was then determined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002752_10775463211010525-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002752_10775463211010525-Figure2-1.png",
        "caption": "Figure 2. Dynamically equivalent manipulator Liang et al. (1998).",
        "texts": [
            " This model because of the momentum conservation law for FFMs has nonlinear parameterization property thus is not feasible for an adaptive control scheme using a linear parameterization model and also does not necessarily have the dynamical properties of a conventional fixed base manipulator. Because our proposed adaptive control approach does not need linear parameterization, we only take advantage of the dynamical properties of conventional fixed base manipulator in the controller design. As shown in Liang et al. (1998), a FFM can be mapped to a conventional fixed base manipulator in which the first joint is a passive spherical one and it is located at the FFM\u2019s center of mass. Also it is assumed that the DEM operates in absence of gravity. Figure 2 shows the equivalent manipulator. In this figure, the orientation of DEM\u2019s first passive joint is represented by (\u03a6b, \u03b8b,\u03a8b), Ji is the joint connecting DEM\u2019s (i 1)th link and ith link, \u03b8i is the rotation of the DEM\u2019s ith link around joint Ji, DEM\u2019s ith link\u2019s center of mass is shown by Ci, ui is the axis of rotation of Ji, lci is the vector connecting Ji to Ci, and Wi is the vector connecting Ci to Ji+1. Again here with calculating DEM\u2019s total kinetic energy and considering q \u00bc \u00bd\u03a6b \u03b8b \u03a8 b \u03b82 \u2026 \u03b8n\u00fe1 T as the vector of generalized coordinates by using Lagrange\u2019s equation, the dynamics of the DEM can be written as M\u00f0q\u00de\u20acq\u00fe C\u00f0q, _q\u00de _q \u00bc \u03c4 (14) where \u03c4 \u00bc \u00bd 0 0 0 \u03c42 \u2026 \u03c4n\u00fe1 is the torque vector of DEM joints, M\u00f0q\u00de 2R \u00f0n\u00fe3\u00de\u00d7\u00f0n\u00fe3\u00de is the inertia matrix, and C\u00f0q, _q\u00de _q2R n\u00fe3 is the vector of Coriolis and centrifugal forces of DEM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001349_10426914.2020.1802035-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001349_10426914.2020.1802035-Figure1-1.png",
        "caption": "Figure 1. Information of studied components and measuring locations.",
        "texts": [
            " Moreover, the influence of casting, natural aging treatment, artificial aging treatment, vibration aging treatment, machining on the residual stress of gray cast iron are analyzed in Section Result and discussion. Through a two-year residual stress tracking of the guideway mounting surface, we found the release law of the residual stress and its impact on the guideway mounting surface precision. Finally, this article is summarized in Section Conclusion. Four types of heavy machine tool components have been studied: Bed, Column, Slide and Headstock. Measuring locations for each component types are schematically illustrated in Fig. 1. The specifications of machine tools examined are concluded in Table 1. Specific physical properties and chemical composition of studied gray cast iron are shown in Table 2. The blind-hole detection method is used in this paper to detect residual stress. Firstly, we attached strain gauge at the measurement point on the workpiece. After that, to release residual stress, we drilled a hole of 2 mm at the strain gauge center. The strain gauge recorded the change in the state of strain once the hole had been drilled (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002511_012004-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002511_012004-Figure9-1.png",
        "caption": "Figure 9. Common shape of hole in gear structure.",
        "texts": [
            " We will proceed with keyhole cutting and disk creation in the Space Claim environment. A design calculation is performed with the model with 60% material removal (Figure 8). This is obtained after the material particles are removed and holes of the corresponding hole or groove shape are cut, as long as the fabrication conditions are fulfilled and aesthetic factors are guaranteed. It is then possible to replace the holes with corresponding large and small round holes, which is very convenient in machining, and optimal shape (Figure 9) is ensured. Based on the results of the standard material density distribution in the above gear models, we will opt for the appropriate gear model option which is close to the optimal result (Figure 19). However, the gears of the given shape are intended to guarantee optimal geometry only, and may have nontechnologic shapes despite satisfying the geometrical optimization. Hence, in the given shape, the form which can be fabricated will be chosen for utilization. Contact stress test: The permissible contact stress is determined (2) with a value of 391",
            " To ensure durable conditions, the force is put at in the most dangerous position on the teeth (right at the top of the teeth). When two gears in mesh, only 1 or 2 pair teeth are touching at a time. The value of the force acting on the tooth is equal to the value of the force during optimal design process. In fact, for large gears (diameter > 400mm), 4-6 round holes are usually cut. Besides, in addition to the common round whole profile, other profiles are also cut, such as alternating grooves, large and small round holes (Figure 9). However, the number of holes must be an even number so as to avoid the eccentricity when the gears are operating. RCMEManuE 2020 IOP Conf. Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10.1088/1757-899X/1109/1/012004 Based on the element removal results in Section 2.1 and practical experience, the researcher cuts the 3D gear model at the disk part with different shapes and rechecks the maximum stress (Figure 10). RCMEManuE 2020 IOP Conf. Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001061_s12652-020-02038-3-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001061_s12652-020-02038-3-Figure3-1.png",
        "caption": "Fig. 3 Hole-area with one, two and three intersections",
        "texts": [
            " Since RE > RS , the size of an EC can be much larger than the subsequent coverage-hole and that is why IEC is used for calculating the hole-area much accurately. Das and Kanti DebBarma (2018) proposed the holearea estimation with one, two or three intersections using computational geometry based approach which is discussed briefly as follows: Let us assume the three sides of the Delaunay Triangle are: a, b and c respectively and there are three situations with one, two and three intersections as shown in Fig.\u00a0 3. Now, area of triangle, At = 1 4 \u221a (a + b + c)(b + c \u2212 a)(c + a \u2212 b)(a + b \u2212 c) , where s = (a+b+c) 2 Figure\u00a03a shows area-estimation with one intersection. Area of hole, Ah = At \u2212 (A1 + A2 + A3) \u2212 (S1 \u2229 S2) where A1 , A2 and A3 are the area of each circular sector with respect to the sensor nodes S1 , S2 and S3. 1 3 Again, A1 + A2 + A3 = R2 S where RS is the sensing radius of each sensor node by assuming Ra = Rb = RS (since each sensor have a fixed sensing radius of RS ). Therefore, with one intersection, area of CH, where d is the distance between two sensor nodes for one intersection. Similarly, Fig.\u00a03b depicts area-estimation with two intersections. Here area of CH, where d1 and d2 are the distances between nodes for two intersections. And finally, area-estimation with three intersections is shown in Fig.\u00a03c, when area of CH, Ah = 1 4 \u221a (a + b + c)(b + c \u2212 a)(c + a \u2212 b)(a + b \u2212 c) \u2212 R2 S \u2212 2R2 S \u22c5 cos\u22121 \ufffd d 2RS \ufffd \u2212 d 2 \u22c5 \ufffd 4R2 S \u2212 d2 Ah = 1 4 \u221a (a + b + c)(b + c \u2212 a)(c + a \u2212 b)(a + b \u2212 c) \u2212 R2 S \u2212 2R2 S \u22c5 \ufffd cos\u22121 \ufffd d1 2RS \ufffd + cos\u22121 \ufffd d2 2RS \ufffd\ufffd \u2212 1 2 \u22c5 \ufffd d1 \ufffd 4R2 S \u2212 d2 1 + d2 \ufffd 4R2 S \u2212 d2 2 \ufffd A h = 1 4 \u221a (a + b + c)(b + c \u2212 a)(c + a \u2212 b)(a + b \u2212 c) \u2212 R 2 S \u2212 2R 2 S \u22c5 \ufffd cos \u2212 1 \ufffd d1 2R S \ufffd + cos \u2212 1 \ufffd d2 2R S \ufffd + cos \u2212 1 \ufffd d3 2R S \ufffd\ufffd \u2212 1 2 \u22c5 \ufffd d1 \ufffd 4R2 S \u2212 d 2 1 + d2 \ufffd 4R2 S \u2212 d 2 2 + d3 \ufffd 4R2 S \u2212 d 2 3 \ufffd where d1 and d2 and d3 are the distances between nodes for three intersections"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure14-1.png",
        "caption": "Fig. 14. On-load magnetic field distribution of CPM-II with \u03b4=90deg. (a) 10A, (b) 12A, and (c) 15A.",
        "texts": [],
        "surrounding_texts": [
            "The back-EMF and electromagnetic torque are shown in Fig. 15. It can be observed that although the PM volumes of CPM-I and CPM-II are similar, the electromagnetic performance of CPM-II is much better than that of CPM-I. In addition, the electromagnetic performance firstly increases and then decreases as \u03b4 increases. When \u03b4=60 deg, it reaches the optimal value. This is consistent with the variation of the air-gap flux density as shown in Fig. 10. It can be found that when \u03b4\u226575 deg, the saturation of the salient iron pole can be eliminated. However, when \u03b4\uff1e90deg, the electromagnetic performance is reduced significantly. Hence, considering the saturation of salient iron pole and the electromagnetic performance, an excellent balance between reduced saturation and good electromagnetic performance can be achieved when 75deg\u2264\u03b4\u226490 deg is satisfied. Considering the manufacturability, the machine with \u03b4=90 deg is chosen for the prototype machine."
        ]
    },
    {
        "image_filename": "designv11_63_0001369_aim43001.2020.9158850-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001369_aim43001.2020.9158850-Figure1-1.png",
        "caption": "Figure 1. A picture of the gripper: (a) two-fingered mode; (b) three-fingered mode. Red box of zoomed-in picture shows the potentiometer.",
        "texts": [
            " (Corresponding author: Pei-Chun Lin: phone: +886-2-3366-9747; fax: +886-2-3366-9914; e-mail: peichunlin@ntu.edu.tw). 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 601 Authorized licensed use limited to: Cornell University Library. Downloaded on August 29,2020 at 07:12:19 UTC from IEEE Xplore. Restrictions apply. presents and discusses the experimental results, and Section VI concludes the work. With advances in technology, the gripper has been developed to become more versatile and function similar to the human palm. In this paper, a dexterous gripper, shown in Figure 1, is developed from our laboratory and utilized to perform the intricate picking up task. The gripper has three fingers; each finger has one active DOF and one passive DOF to create compliant adaptability, because higher adaptability means more stability as the gripper executes a task. In addition, the gripper can change its finger configuration to generate two modes: two-finger mode and three-finger mode. This gripper leverages a hall sensor as our proximity sensor to determine which mode configuration the gripper should adopt, and within the palm of the gripper a servo motor is implanted to change the mode",
            " In general, objects can be classified into some primitive shapes, such as circle, rectangle, etc. In this work, the covariance calculated from principal component analysis (PCA) is utilized as criteria to distinguish a symmetric object from an asymmetric object, and this will be discussed further in a later section. Using these classifications and finger configuration modes, the corresponding grasp types are listed in Table I. To avoid a collision with other objects while the gripper carries out the work, this gripper has been set up with a potentiometer on each finger, as shown in Figure 1, to detect the rotating angle of the finger. An algorithm is proposed to calculate the width of the object, and position control is used to adjust the finger to a particular angle. In section IV, a detailed explanation of the algorithm will be provided. In this section, a method will be deliberated for how to collect data and train Q-PointNet through the steps mentioned in Section III A, B, and C. First, Mask R-CNN will be introduced, which is developed to generate a mask from the targeted object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001845_s00202-020-01132-1-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001845_s00202-020-01132-1-Figure1-1.png",
        "caption": "Fig. 1 Different housing structures of TEFC machines [3, 4]",
        "texts": [
            " The proper cooling method of an electrical machine is selected according to several different parameters, e.g., machine application, machine topology, machine current density, speed, the operational environment, cost, and machine size [1, 2]. The conventional cooling system implemented in industrial electricalmotors is a totally enclosed fan cooled (TEFC). In this cooling system, the external fan is attached to the non-driven part of the machine to blow the fluid toward the semi-open fin channels (Fig. 1) [5]. In this approach, the fin channels over the housing are essential [1]. B Payam Shams Ghahfarokhi payam.shams@taltech.ee 1 Department of Electrical Machines and Devices, Riga Technical University, Kal\u0327k\u0327u iela 1, Riga, Latvia 2 Department of Electrical Power Engineering and Mechatronics, Tallinn University of Technology, Ehitajate tee 5, Tallinn, Estonia 3 Department of Electrical Engineering and Automation, Aalto University, Espoo, Finland One of the most critical parts of the thermal analysis of TEFC machines is determining the heat transfer from the machine housing to the ambient"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001727_s0263574720000843-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001727_s0263574720000843-Figure1-1.png",
        "caption": "Fig. 1. The 3-3-PUS parallel mechanism.",
        "texts": [
            "14 The goal of this study is to further diminish the positioning error to improve the kinematic accuracy of the docking simulator. The content of this paper is listed as follows. Section 2 presents the kinematic model of the 3-3-PUS parallel robot. Section 3 studies the kinematic parameters of the robot and shows how the parameters separation is conducted. Then, a simplified error model is derived. In Section 4, the calibration experiment and the confirmatory experiment are conducted. Then, the results and concluding remarks are presented in Section 5. As seen in Fig. 1 this parallel mechanism has one moving platform and six uniform kinematic chains. Each chain is composed of a spherical (S) joint, a linkage, a universal (U) joint, and a prismatic (P) joint in sequence. So each chain is the PUS kinematic chain. The six prismatic joints are linear driving joints, which are installed on the fixed base. The moving platform is the end effector. According to the Gr\u00fcbler\u2013Kutzbach\u2019s criterion for mobility calculation,32 the DOF of this mechanism can be calculated as follows: M = 6 \u00d7 (14 \u2212 18 \u2212 1) + (6 \u00d7 1 + 6 \u00d7 2 + 6 \u00d7 3) = 6 https://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002215_s41693-020-00051-8-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002215_s41693-020-00051-8-Figure3-1.png",
        "caption": "Fig. 3 Robotic Oscillating Wire Cutting setup, with an ABB 6620. The programming of ROWC is carried on in Rhino/GH environment using Robots I/O plugin",
        "texts": [
            " In fact, as the wet clay mix is very dense, the oscillation allows for the creation of a very narrow diagonal trajectory and incorporates an air gap, which allows for the quick cutting of geometries presenting undulations and very tight curvatures that would be very challenging to achieve with a static wire due to the \u201cdrag\u201d phenomena. The oscillation of the wire helps deal with the resistance of cutting through the material and although non-static, provides a swift, smooth cut without geometric distortion. The ROWC end effector is custom made and proprietarily designed by the authors. An electrical engine propels a structural frame holding a suspended wire of 1.5\u00a0mm diameter in a reciprocating motion. The wire is tensioned across the frame manually using bolts. The cutting is executed on a vacuum table (Fig.\u00a03) which provided enough stability, combined with the weight of the wet clay, for the shallow brick designs, for the longer profiles stabilizing pressure from top was needed. Comparative tests were performed for ROWC vs. static wire (RSWC) end effectors on 7 shape variant sample designs cut from a 55 \u00d7 205 \u00d7 105 mm wet clay block base. 1 3 At identical suspension pull and wire thickness, RSWC cutting failed on 5 of 7 profile designs, while resulting in significant Tool Centre Point deflection on the remaining 2, while cutting at fastest 1\u00a0mm/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002138_s11668-020-01106-2-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002138_s11668-020-01106-2-Figure3-1.png",
        "caption": "Fig. 3 Equivalent stress cloud diagram of rollers with different filling degrees",
        "texts": [
            " Among them, the composite structure has the highest degree of improvement to the bending stress, followed by the equivalent stress, and the contact stress is the lowest. Equivalent Stress Analysis The filling degree (k) of the composite cylindrical roller is the ratio of the minimum hollow diameter of the roller shell to the diameter of the roller. To study the effect of filling degree on composite cylindrical rollers, numerical calculation models for composite cylindrical rollers with filling degrees of 40, 45, 50, 55, 60, 65 and 70% were established, respectively. Both apply a load of 12 kN. Figure 3 is the equivalent stress cloud diagram of composite cylindrical rollers with different filling degrees. It can be seen from the figure that the equivalent stress on the composite cylindrical roller is mainly concentrated on the contact area between the roller and the inner and outer rings of the bearing and the corresponding inner wall. There is obvious stress concentration on the two ends of the contact surface between the roller and the inner and outer rings of the bearing. With the increase in the filling degree, the stress concentration on the contact surface becomes less obvious"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002630_iccia52082.2021.9403540-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002630_iccia52082.2021.9403540-Figure1-1.png",
        "caption": "Figure 1. Mobile robot in configuration space",
        "texts": [
            " Afterward, the idea of introducing tubes in LMPC is for achieving robust property to modeling uncertainty is proposed in Section 4. Simulations are presented in Section 5. II. MODELING This section presents the mathematical modeling of a mobile robot under slip conditions for trajectory tracking. The kinematic modeling of a mobile robot is defined as follows: ( )cos ; ( )sin ; ( ) 2 2 l r l r r lx r y r r h \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03b8 \u03b8 \u03b8 + + \u2212 = = =&& & (1) where, and denote the angular velocity of the left and right wheels, and h is the distance between the center of the wheels as it showed Figure 1. According to the road condition, there is a difference between the actual and theoretical linear velocity of wheels. Therefore, the slip is formulated as the following 1 v i r\u03c9 = \u2212 (2) Where 0 \u2264 i < 1 and v and r denote the wheel's actual linear velocity and the wheel radius, respectively. It is assumed while moving at relatively low speeds, the dynamic effects on a robot and lateral slip effects are neglectable [18]. Therefore, by incorporating longitudinal slip effect on mobile robots, it is modeled by: 1 1 0 2 2 0 1 1 0 1 l r l l r r i i x cos r y sin i i r h h \u03b8 \u03c9 \u03b8 \u03c9 \u03b8 & & & \u2212 \u2212           =        \u2212 \u2212     \u2212       (3) A reference trajectory in the global frame is adopted to control the mobile robot to track a predefined trajectory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure2-1.png",
        "caption": "Fig. 2. Influence of inclined links on workspace/boundary singularities.",
        "texts": [
            " San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at In practice, the load-bearing capability can be significantly improved with inclined links. It is also obvious that non-planar links are going to influence many other characteristics. For example, increasing the inclination angle improves the load-bearing capability and stiffness, but it may affect the workspace or Type I singularities. To drive home this point, a single link with a rotary joint R is illustrated (Fig. 2). The link shown in black color is placed in the XY plane. Another link in green is of the same size (L1) and is inclined (the method of inclination is described in Fig. 3). Both the links are rotated around the Z axis. The planar link covers a work area C1, whereas the inclined (spatial) link covers C2. It will be of the readers\u2019 interest to know the appropriate inclination angle which does not compromise any significant feature of the manipulator. The following features of the manipulator, such as the moment of inertia (MOI), stiffness, workspace, inertia forces and the moving mass of the manipulator are addressed in this article"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002142_jmems.2020.3047774-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002142_jmems.2020.3047774-Figure2-1.png",
        "caption": "Fig. 2. Geometric model of hinge deformation.",
        "texts": [
            " Downloaded on May 17,2021 at 21:08:49 UTC from IEEE Xplore. Restrictions apply. sensor integrated with the origami sheet device and evaluate the characteristics of the SWCNTs strain sensor. Then, we demonstrate that our origami sheet device can measure the shape of an object with recognizing the difference of the bend and the stretch separately. To design the shape-recognizable origami sheet device, the relationship between the bending angle of a hinge, \u03b8 , and the strain of the SWCNTs layer, \u03b5, is expressed by a simple geometric model (Fig. 2). We assume that the deformation of the hinges where SWCNTs strain sensors are formed on the surface is always an arc when the hinge is bent. As the bending angle \u03b8 increases, the radius of curvature, r , decreases and the strain on the SWCNTs layer increases. The relationship between the bending angle \u03b8 (degree) and the strain \u03b5, is represented below, \u03b5 = \u03c0\u03b8 t 360L Hinge ( = r + t 2 r ) (1) where the length of the hinge is L Hinge and the thickness of the hinge is t . As shown in equation (1), the strain varies linearly with the bending angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000733_lra.2020.2979633-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000733_lra.2020.2979633-Figure12-1.png",
        "caption": "Fig. 12. Experimental setup of stiffness-controlled hopping leg.",
        "texts": [
            " Table II shows the changes of the apex heights during hopping, which shows the apex heights of hopping trajectories in Fig. 10 and Fig. 11. The reduction rates of two consecutive apex heights are also described in Table II, which shows that the reduction rate of the apex height can be minimized by the proposed optimal landing condition. The experimental studies carried out to confirm the effectiveness of the proposed method. The overall experimental setup with SEA-driven robotic leg is shown in Fig. 12. The robot leg is stiffness-controlled and is laid on a force plate which measures the ground reaction force. The optical tracker is set to track the markers which are attached to the hip, the knee, and the end-effector to measure the trajectories of the body and the foot. The trunk is pushed to the ground by a hand, by which the initial condition is set that can coincide with the proposed condition. Figs. 13(a) and (b) are the results of the hopping experiment; Fig. 13(a) is the case where initial condition is set satisfying the proposed optimal landing condition as shown in the magnified plot in the middle figure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002004_ddcls49620.2020.9275203-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002004_ddcls49620.2020.9275203-Figure1-1.png",
        "caption": "Fig. 1: Hummingbird",
        "texts": [
            " Besides, MFAC has been widely applied to many practical industrial processes such as distillation system, linear motor, injection modeling process, subway trains, autonomous car, and pH neutralization process [14\u201320]. During the flight of the aircraft, the lift force and thrust are mainly generated by the wings. Aiming at the difficulty of controller design caused by the nonlinearity, time-varying and strong coupling of FWMAV, a thin tail is added to the tail as Ref. [21]. Then it is used to control the pitch angle independently. The overall model is similar to a hummingbird, as shown in the Figure 1. Normally, the actuator is a vibration motor with a maximum frequency, which is the threshold of the control input. Moreover, input saturation makes the output value of the system unable to meet the reference value, which may even lead to the instability of the system. 978-1-7281-5922-5/20/$31.00 \u00a92020 IEEE DDCLS'20 627 Authorized licensed use limited to: Carleton University. Downloaded on June 03,2021 at 03:40:35 UTC from IEEE Xplore. Restrictions apply. The paper is organized in the following manner: Section II introduces the dynamic model of the vehicle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002220_icpai51961.2020.00020-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002220_icpai51961.2020.00020-Figure3-1.png",
        "caption": "Fig. 3: Example of a trajectory problems.",
        "texts": [
            " The accuracy of shuttlecock detection is the major factor in the construction of badminton trajectory and directly affects the performance of shots detection. However, due to the high moving speed, the badminton images in the video may be blur and cannot be detected by TrackNet. According to the statistics, the shuttlecock trajectory can be depicted from broadcast video with an accuracy of around 80%. Therefore, in this paper, we proposed the trajectory smoothing algorithm to improve the accuracy of TrackNet. TrackNet suffers from two major problems which may affect the performance of shots detection. Figure 3(a) refers to the normal trajectory, the trajectory should approximate to the curve of second order and coordinates of each frame should exist. However, video conditions like badminton images are blur or background noise which looks like badminton will make TrackNet misjudge accurate coordinates in the video. When the Euclidean distance of predicted coordinates and actual coordinates is larger than 10 pixels, we defined this situation as shifted coordinate, as shown in Figure 3(b). Another situation like coordinates are empty because of TrackNet can not locate badminton in the video, we defined this situation as missing coordinates, as shown in Figure 3(c). In order to solve these problems, the proposed algorithm can be separated into three phases to discuss. The first phase is denoise. In this phase, we want to eliminate shifted coordinates as much as possible and leave the accurate coordinates for the next phase. According to the statistics, around 9% of predicted coordinates shifted more than 100 pixels and 2.4% of predicted coordinates shifted between 10 pixels and 100 pixels. Therefore, in this phase, we used the Euclidean distance between coordinates of each frame to remove the predicted coordinates which shifted more than 100 pixels",
            " Predicted coordinates shifted between 66 Authorized licensed use limited to: Walailak University provided by UniNet. Downloaded on May 17,2021 at 04:13:44 UTC from IEEE Xplore. Restrictions apply. 10 pixels and 100 pixels can not be removed because of the distance between abnormal coordinates is no larger than the distance between normal coordinates. Thus, we will deal with it in the next phase. The second phase of this algorithm is curve fitting. In this phase, we used quadratic equations which composed of coordinates from the first phase to smooth and fill the trajectory. As shown in Figure 3(d), We define seven consecutive sets of coordinates as check window. If there existed at least three coordinates in this check window, we can calculate quadratic equation of these coordinates. According to this quadratic equation, we defined the minimum distance from coordinate before the check window to the curve as front check distance. Similarly, the minimum distance from coordinate after the check window to the curve as back check distance. For each coordinate, it has the back check distance of the previous check window and the front check distance of the following check window. If the back check distance and the front check distance are abnormally large, it means that the current coordinate deviates the trajectory which composed of 1check windows nearby it. Therefore, this coordinate will be defined as shifted coordinate, as shown in Figure 3(e). With the help of this algorithm, we can remove shifted coordinates which phase one can not detect. On the other hand, if the back check distance and the front check distance are smaller than 5 pixels, it means that current coordinates fit the trajectory which composed of check windows nearby it. Consequently, we defined this coordinate as a match point, as shown in Figure 3(f). In this phase, we used match points as the filling standard. When the match point happened and there are missing coordinates or shifted coordinates in the check windows nearby it, we can use the quadratic equation of each check windows to fill coordinates. After the second phase, most of the coordinates had been adjusted and the trajectory becomes more complete. However, there are still some missing coordinates left behind. Thus, in the third phase, we used interpolation to fill the rest of the missing coordinates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000312_1464419319889158-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000312_1464419319889158-Figure2-1.png",
        "caption": "Figure 2. Dynamic model of sun\u2013planet ith pair.",
        "texts": [
            " Thus, the quality, support stiffness, and other parameters are the same. The left and right sides of the herringbone gears are ideally symmetrical. In the following derivations, the subscripts s, p, c, and r in the relevant parameters represent the sun gear, star gear, planetary cage, and ring gear, respectively. The subscripts L and R represent the left and right sides of the herringbone gear, respectively. Because of the large number of model components, the finite element method was used to model the system\u2019s dynamics.7 Figure 2 shows the mesh of the ith star gear and sun gear, in which the installation angle of the ith star gear is pi \u00bc 2 i 1\u00f0 \u00de=N. The angle between the plane of the meshing line of the gear pair and the y-axis of the center component is spi. The expression for spi is given in equation (1), wherein is the end pressure angle of the sun gear. The mesh dynamic equations for one side, taking the right side of the sun gear as an example, are given in equations (2) and (3). Because the friction coefficient has little effect on the vibration mode, the friction and lubrication are not taken into account in equations (2) and (3) spi \u00bc s pi \u00f0counterclockwise rotation of sungear\u00de pi s \u00f0clockwise rotation of sun gear\u00de 8>>>< >>>: \u00f01\u00de ms \u20acxsR\u00f0t\u00de \u00fe kspn cos sin spnpspnR \u00bc 0 ms \u20acysR\u00f0t\u00de \u00fe kspn cos cos spnpspnR \u00bc 0 ms \u20aczsR\u00f0t\u00de kspn sin pspnR \u00bc 0 Js \u20ac sR\u00f0t\u00de \u00fe kspnrs cos pspnR \u00bc 0 8>>>< >>>: \u00f02\u00de mpi \u20acxpiR\u00f0t\u00de kspi cos b sin tpspiR\u00f0t\u00de \u00bc 0 mpi \u20acypiR\u00f0t\u00de kspi cos b cos tpspiR\u00f0t\u00de \u00bc 0 mpi \u20aczpiR\u00f0t\u00de \u00fe kspi sin bpspiR\u00f0t\u00de \u00bc 0 Jpi \u20ac zpiR\u00f0t\u00de\u00fekspirp cos bpspiR\u00f0t\u00de \u00bc 0 8>>>< >>>: \u00f03\u00de where m is the unilateral mass of the herringbone gear, J is the moment of inertia around the z-axis, and r is the radius of the base circle ( \u00bc s, pi)",
            " The quantity pspiR\u00f0t\u00de is the relative displacement of the right side of the herringbone gear along the direction of the meshing line of the engaging teeth pspiR\u00f0t\u00de \u00bc \u00bdxsR sin spi \u00fe ysR cos spi xpiR sin s ypiR cos s \u00fe rs zR \u00fe rp piR cos b \u00fe zpiR zsR sin b Rearranging equations (2) and (3) yields Ms 0 0 Mp \u20acqsR\u00f0t\u00de \u20acqpiR\u00f0t\u00de \" # \u00fe K1 spi K2 spi syms K3 spi \" # qsR\u00f0t\u00de qpiR\u00f0t\u00de \u00bc 0 0 By analyzing the meshing of one side of the internal gear and the ith star gear, we can also obtain expressions for K1 rpi, K 2 rpi, K 3 rpi, K 1 cpi, K 2 cpi, and K3 cpi. The coupling matrices for the left and right sides of the herringbone gear are derived below. Taking the sun gear as an example, as shown in Figure 2, the combined nondamped motion equations for the left and right sides of the sun gear are ms \u20acxsL\u00f0t\u00de \u00fe kssx\u00f0xsL\u00f0t\u00de xsR\u00f0t\u00de\u00de \u00bc 0 ms \u20acysL\u00f0t\u00de \u00fe kssy\u00f0 ysL\u00f0t\u00de ysR\u00f0t\u00de\u00de \u00bc 0 ms \u20aczsL\u00f0t\u00de \u00fe kssz\u00f0zsL\u00f0t\u00de zsR\u00f0t\u00de\u00de \u00bc 0 Js \u20ac zsL\u00f0t\u00de \u00fe kss z\u00f0 zsL\u00f0t\u00de zsR\u00f0t\u00de\u00de \u00bc 0 8>>>< >>>: \u00f04\u00de ms \u20acxsR\u00f0t\u00de \u00fe kssx\u00f0xsR\u00f0t\u00de xsL\u00f0t\u00de\u00de \u00bc 0 ms \u20acysR\u00f0t\u00de \u00fe kssy\u00f0 ysR\u00f0t\u00de ysL\u00f0t\u00de\u00de \u00bc 0 ms \u20aczsR\u00f0t\u00de \u00fe kssz\u00f0zsR\u00f0t\u00de zsL\u00f0t\u00de\u00de \u00bc 0 Js \u20ac zsR\u00f0t\u00de \u00fe kss z\u00f0 zsR\u00f0t\u00de zsL\u00f0t\u00de\u00de \u00bc 0 8>>>< >>>: \u00f05\u00de Here, kssx, kssy, kssz, and kss z represent the radial and axial translation and rotational stiffness values of the shaft section between the left and right rows of teeth in the sun gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000412_phm-qingdao46334.2019.8942851-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000412_phm-qingdao46334.2019.8942851-Figure1-1.png",
        "caption": "Fig. 1. Rendered perspective view (top) and X-ray front view (bottom) of 3-D printed planar array of 49 internally fed electrospray emitters. Device features include tapered internal channels, threaded holes, and spill protection features.",
        "texts": [
            " The design of the 3-D printed electrospray sources is described in Section II, while the fabrication of the devices and the exploration of the resolution capabilities of the fabrication technique are documented in Section III. The experimental characterization of the liquid ionizers is described in Section IV, and a discussion of the results is provided in Section V. Perspective and front views of a computer-aided design (CAD) file that describes a 3-D printed, multiplexed electrospray source are shown in Figure 1. The device is a planar array of miniaturized, internally fed electrospray emitters. The main frame is a solid block 3 mm thick with a square base 24 mm wide. The frame includes 4 evenly spaced 4-40 threaded through-holes that can be used to clamp the device to a chuck with a liquid feedthrough. The design of the device includes the following features: - Each emitter is internally fed by a slightly tapered converging channel with the intention to decouple the pressure required to fill-in the emitter and the pressure needed to set a flow rate through the emitter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure4-1.png",
        "caption": "Fig. 4. CAD model of the mechatronic assembly, with transparent nonconductive bed.",
        "texts": [
            " All these considerations are factored under the assumption that the conductive traces are designed at the limit of breaking if a tighter fit would be attempted, thus allowing the maximum possible current to be driven by the battery through the conductive 3D printed parts and into the motor. A preliminary design was created in NX 12 [21] for both conductive and non-conductive elements. We have chosen to ensure that the traces can form a shoulder resting on the battery main package, as can be seen in Figure 3. This design aspect is convenient since it allows for the conductive paths to be evenly clamped between the non-conductive holder and the battery main package, until exposing themselves to the motor terminals, as visible in Figure 4. This reduces the risk of undesired rotation and misalignment of the traces around the battery terminals. The non-conductive holder consists of two elements, housing the battery with conductive traces and the motor, respectively. In order to guide the battery and help maintain its position within the mechatronic assembly, the bed is dimensioned such that the length of the bed at the base of the wedge is 0.5 mm shorter than the battery, giving an approximate strain of 1%, within the usual elongation range of PLA [22]. The selected standard component is a low-current Tamyia RC300-FT-14270 DC motor [23], which is held in place by the wedge system described previously. The position of battery and motor within the 3D printed mechatronic assembly is illustrated in Figure 4. The principle of holding the motor in place may be expanded to virtually any other DC motor. Previous efforts have seen a different topology of motor being successfully wedged into a 3D printed mechatronic assembly [7]. Printing the mechatronic assembly as a united entity using dual extruders has raised the challenge of inserting the battery within the conductive traces without the risk of damaging them. Due to these relatively high accuracy assembly requirements for dual extruder printing, one can also resort to printing the conductive and non-conductive parts separately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000555_j.jmapro.2020.01.054-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000555_j.jmapro.2020.01.054-Figure4-1.png",
        "caption": "Fig. 4. Schematics of the models and actual pictures of the components of the three different forming schemes: models in (a) BO-3, (b) PL, and (c) GR-3; components in (d) BO-3, (e) PL, and (f) GR-3.",
        "texts": [
            " The substrate was forged at a temperature of 40 \u00b0C below the phase transformation point (998 \u00b0C) and air-cooled to room temperature; the deformation amount was controlled within 20 %\u201340 %. The subsequent heat treatment process involved heating at 780 \u00b0C for 40min, holding for 2 h, and then air-cooling. The powder was produced by the rotating electrode method, where the particle size is in the range of 75\u2013150 \u03bcm, as shown in Fig. 3. The present study proposed three schemes of composite forming, i.e., depositing single-pass multilayer walls by LSF on Ti-6Al-4 V forged substrates with three different geometries: with a groove, platform, and boss. Fig. 4 shows the schematics and photographs of the three components, where the blue-colored part stands for the substrate and the red one for the AM part. The AM part is deposited along the longitudinal edge of the substrate. Hereinafter, the schemes for forming the components with a boss, platform, and groove are referred to as BO-3, PL, and GR-3, respectively. One noteworthy point is that different from the PL component, those in BO-3 and GR-3 had a 3mm high boss and a deep groove on the edge of the substrate, respectively. The dimensions of all substrates in the different schemes were 100mm\u00d720mm\u00d730 mm (see Fig. 4 for more details). To avoid blocking the output path and stroke of the powder nozzle and the laser head, the inverted trapezoid groove which was wide at the top and narrow at the bottom was designed in GR-3, and all three sides of the groove were processed into inclined ones with an angle of 45\u00b0. As a result, the deposited length on the groove was<64mm, while the deposited length in other areas of the component in GR-3 or in the other two components was 64mm. The AM part was deposited using a 2mm diameter laser beam by reciprocating scanning with simultaneous powder feeding. However, due to the existence of the groove and boss, the deposited heights of the three components were different, i.e., 18mm, 15mm, and 12mm for GR-3, PL, and BO-3, respectively, and it was ensured that the final components have the same height. The coordinate system used throughout this study is shown in Fig. 4. The X, Y, and Z axes represent the longitudinal, transverse, and normal directions, respectively, and the+X direction is the laser scanning direction of the starting layer. Point O in the bottom left vertices of the substrate was set as the origin. The LSF machine fabricating the samples is BLT-C600 (Bright Laser Technologies Co., Ltd., Xi'an, China), as shown in Fig. 5. All three components were deposited using a fiber laser in an argon atmosphere, using the same material and process parameters",
            "5\u00d75 mm2, and the range of the azimuth angle \u03a8 of the diffraction crystal plane was from \u22125\u00b0 to 45\u00b0 (10 points). The measured diffractive crystal plane was the (213) hcp plane, and its diffraction angle 2\u03b8 was 139.69\u00b0. The Gaussian fitting method was used to locate the peak and corrections were made for the background, absorption factor, and Lorentz-polarization factor. The elastic modulus E and the Poisson\u2019s ratio \u03bd were 110 GPa and 0.31, respectively. According to the shape and size of the actual components (see Fig. 4) in the above three schemes, three coupled thermomechanical models were established in the ANSYS finite element software to simulate the deposition process of the single-pass multilayer walls. The finite element mesh of the components is shown in Fig. 8. The numbers of layers of the models in GR-3, PL, and BO-3 were 180, 150, and 120, respectively. Except for the area near the groove in GR-3 that used a free mesh due to the complex structure, other areas of the three models used a mapped mesh. To improve the accuracy of calculation, a finer mesh was used at the location close to the cladding because the thermal and stress gradients at that location are usually the largest"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001041_iet-epa.2020.0237-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001041_iet-epa.2020.0237-Figure4-1.png",
        "caption": "Fig. 4 Equivalent insulation method in the stator slot [23]",
        "texts": [
            " The equivalent thermal conductivity of the slot can be evaluated by three methods: layer winding, cuboidal, and equivalent insulation. In this research paper, the equivalent insulation method is applied to find the equivalent conductivity of the slot. This method is a fast calculation method and can be easily used by knowing the slot-filling factor. Based on this method, copper has high thermal conductivity and low conduction thermal resistance. Consequently, the copper thermal resistance can be neglected in the analytical model of the stator winding in the circumferential and radial direction. Fig. 4 shows a simplified assumption to model the equivalent insulation method in the active part of the slot. As a result, the equivalent thermal conductivity of the insulation kins is assessed by a linear empirical correlation as in [22] kins = 0.1076k f + 0.029967. (9) Table 1 Description of thermal resistances Component Description R10 convection resistance of the air gap R11 conduction resistance of the upper half side of the rotor R12 conduction resistance of the lower half part of the rotor R13 axial conduction of the shaft Pir iron losses in the stator PJa active copper losses PJe end-winding copper losses Peddy iron losses in the rotor Table 2 Geometrical data Name Symbol Unit Value stator core length Ls mm 156 stator inner diameter Dis mm 136 stator outer diameter Dos mm 219 number of slots Ns \u2014 36 air-gap height lg mm 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure48.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure48.1-1.png",
        "caption": "Fig. 48.1 a K-wire guide dimensions and section view, b FEA of guide, c inclined orientation of k-wire guide (good design), d k-wire obstructing the tool path form above, e parallel orientation of k-wire guide (bad design) and f from below",
        "texts": [
            " The FEAwas carried usingABSM30i asmaterial for the k-wire guide. The result of equivalent stress shows that the design is sufficiently strong, and the stresses are well within limits. The guide should have sufficient height to prevent misorientation. The height of the guide k-wire guide is fixed to 10 mm as observed by actual 3D printing and inserting k-wire. The k-wire must not be placed too closed to each other. Two k-wire must be at least 1 cm apart. If the diameter of k-wire is d, then outer diameter of k-wire guide is taken as d + 5 mm as shown in Fig. 48.1a. Minimum three k-wire is necessary to restrict the moment of the jig. To prevent this, the k-wire must not be parallel to each other. A slight inclination helps to restrict the upward moment of the jig. The minimum angle of 100\u2013150 between two k-wire is sufficient to restrict the lifting of k-wire. Care must be taken that k-wire should not obstruct the cutting tool path. Contact Surface: The contact surface of the jig is the surface that comes in contact with the bone. Bone is covered with a layer of tissue called the periosteum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure6-1.png",
        "caption": "Fig. 6. Combined kinematic singularities: (a) first configuration; (b) second configuration.",
        "texts": [
            " (10) allows for the derivation that the 2UPR-PRU PKM is in the forward kinematic singularity only when limb 1 or 2 is co-linear with 1 2A A . Therefore, two forward kinematic singular configurations occur, as shown in Fig. 5. 3.2.3 Combined kinematic singularity Only when =qJ 0 and =\u03b7J 0 are satisfied simultaneously, the PKM is in the combined kinematic singular configuration. These results show that combined kinematic singularity occurs only when limb 1 or 2 is co-linear with 1 2A A , and limb 3 is simultaneously perpendicular to slider 3OB , as shown in Fig. 6. From Figs. 4-6, one can find that a good choice of link parameters can enlarge the workspace region. 4. Dynamic analysis Dynamic modeling of 2UPR-PRU PKM based on the screw theory and principle of virtual work can be used to explain the mapping between the actuated forces and the output trajectory, which is useful for the development of efficient control system in practical applications. In addition, the dynamic performance evaluation of this PKM is also essential for potential industry applications that need good dynamic response"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001459_978-981-15-4488-0_74-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001459_978-981-15-4488-0_74-Figure3-1.png",
        "caption": "Fig. 3 Schematic diagram showing the standard size of flexural test specimen as per ASTM D790 standard (all dimensions are in mm)",
        "texts": [
            " Flexural properties of FRP materials were evaluated as per ASTM D790 standard test method. The standard dimensions of the specimen for flexural test are 126 mm\u00d7 12.5 mm \u00d7 t mm where t represents the thickness of the specimen. The composite specimens are properly cut to the standard dimensions. The flexural test was conducted on the same INSTRON machine with feed rate of 1.2 mm/min. The hybrid specimen is placed on two knife edges provided on the testing machine with equal over-hanging on both sides. A transverse load is applied at the center point of the specimen. Figure 3 shows the standard size of the composite specimen used for the flexural test. Free vibration characteristics such as modal frequency and modal shapes are evaluated for glass/hemp hybrid composite laminates under fixed-free end condition. The preferred glass/hemp hybrid laminates are cut into the standard size of 250 mm \u00d7 25 mm \u00d7 t mm where t represents the thickness of the specimen for free vibration study. For the effective transformation of excitation force into the composite beam and to achieve fixed-free boundary condition, a fixture is used to clamp the composite specimen"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002570_iros.2011.6048138-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002570_iros.2011.6048138-Figure3-1.png",
        "caption": "Fig. 3. The finite element model of foil type stator. The number of the Node is 90,795, the number of the element is 39,163.",
        "texts": [],
        "surrounding_texts": [
            "Fig.1 shows the structure of the ultrasonic motor with a foil type stator. When ultrasonic vibration is transmitted to the waveguide from the ultrasonic vibrator, the vibration changes to a traveling wave and propagate to the foil type stator through the waveguide. At that moment, friction force is generated between the stator and the rotor by the vibration of the foil type stator. The rotor is rotated by this friction force. The foil-type ultrasonic motor is found to rotate in the same direction as the direction of the traveling wave; this phenomenon does not match the general ultrasonic motor principle. Xie[5] already analyzed characteristics of surface 978-1-61284-456-5/11/$26.00 \u00a92011 IEEE 750 particle motion of the coiled waveguide. However, the analysis was performed for the coil-shape stator with square shape in cross-section. The foil-type stator has plate-shape cross-section; it is not able to compare easily with previous research. In the research field of nondestructive inspection, Lamb wave[6] is well known as the propagation wave in a plate. It propagate in a plate toward the longitudinal direction with flexural or dilatational vibrations and could be propagate in the foil-type stator. But Lamb wave propagate with very complex characteristics, varying with frequency, material properties, and thickness of a plate. And also many modes with different velocities excited at the same time[7], [8], [9]. Because of the Lamb wave complication, it is difficult to solve the stator motion by analytical method. Consequently, in order to investigate the driving principle of the motor, finite element method using ANSYS (ANSYS Inc.) is performed to observe the shape change of the foil type stator. The model for FEM made by 3DCAD is shown in Fig. 2, and the specifications of the model parameters are shown in Tables I, II, III. The analysis model is divided into hexahedral elements with 20 nodal points. The input force from the wave generator is as follows; Fin = 0.000014 \u00b7 sin(2\u03c0 \u00b7 26800 \u00b7 t) (1) This equation is the same as for the wave generator using in a real experimental setting. Fig.5 shows the result of the transient response analysis. First, a longitudinal wave is propagated to the foil stator and then the stator deforms in the radial direction. Deformation of the stator is transmitted from the joint of the stator and waveguide to the tip of the stator. To quantify this deformation, the displacement of the stator surface was observed from the side of the waveguide side. The observation points are shown in Fig.6. Figs. 7, 8, 9 show the analysis result of surface displacement in the y-z coordinate for each turn of the foil. Except for the bottom side, the result shows every observation point rotating in the winding direction of the foil. Lamb wave usually consists of many modes. but in low ultrasound frequency like our experimental situation, it has mainly two modes, symmetric (S0) mode and asymmetric (A0) mode [8]. The surface of the symmetric mode rotates forward direction, and the symmetric mode rotates reverse direction (Fig. 10) . From the experimental result, it is apparent that symmetric wave is dominant mode of the stator."
        ]
    },
    {
        "image_filename": "designv11_63_0003101_s42417-021-00349-z-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003101_s42417-021-00349-z-Figure1-1.png",
        "caption": "Fig. 1 Hybrid bearing assembly",
        "texts": [
            " The main aim of the present work is the coupled numerical model of the rotor, coupling the rotor equations of motion with the fluid film forces from the GFBs and electromagnetic forces from the AMBs. The non-dimensionalization of the magnetic forces is done in line with the fluid film forces and the Reynolds equation. Accordingly, the unbalance responses of rotor supported on GFBs and hybrid GFBs are compared based on the amount of unbalance. The ability to sustain unbalance occurring at an arbitrary time in the rotor-bearing system has been investigated. The assembly of the hybrid GFB in the present investigation is shown in Fig.\u00a01. The top foil and bump foil are placed in between the actuator and the rotor. The top foil and the bump foil has been assumed to be made of non-magnetic material like INCONEL X-750, such that it does not obstruct the path of the magnetic flux. However, due to the incorporation of the top foil and the bump foil, the air gap is increased. Higher air gap requires higher current to produce the required magnetic flux. 1 3 A hetero-polar active magnetic bearing has been used as an electromagnetic actuator (EMA) for the hybrid GFB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000344_iecon.2019.8927327-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000344_iecon.2019.8927327-Figure5-1.png",
        "caption": "Fig. 5. Prototype of 39-slot and 12-pole motor with UBW structure",
        "texts": [
            " In addition, 39-slot and 6-pole machine has the minimum THD level with a sinusoidal fundamental harmonic since this machine has one of the lower pole number. On the other hand, 39-slot and 24-pole machine has the highest THD level. IV. TEST RESULTS OF A 12-POLE UBW MOTOR PROTOTYPE \u2013 A CASE STUDY A prototype motor with unbalanced winding and 39-slots and 12-poles has been manufactured to verify the motor performance results experimentally. The stator and rotor laminations and mounted un-skewed rotor structure are illustrated in Fig. 5. In addition, 2-dimensional (2D) finite element analysis (FEA) 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 1000rpm between FEA and test results is shown in Fig. 6. 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. Moreover, on-load tests are performed, and the torque output of the proposed motor is attained for different load conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001026_ab9549-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001026_ab9549-Figure6-1.png",
        "caption": "Fig. 6. Setup used for actuation analysis of DEA in unimorph configuration.",
        "texts": [
            " Therefore, dielectric properties were measured through dielectric spectroscopy (Novocontrol Alpha-A) in the frequency range of 0-105 Hz under contact electrode mode as schematically shown in Fig. 5. A gold coated electrode having diameter of 20 mm was used in all the tests. The dielectric permittivity and dielectric loss factor of all the samples were recorded at a test voltage of 3 V. The thickness of all the samples was kept constant at 125 \u00b5m. The effect of CNT addition on the tip displacement of bend actuator was compared through actuation test as shown in Fig. 6. The breakdown current was set at 20 \u03bcA and the voltage was applied in the incremental steps of 200 V per 15 seconds. Actuation of DEA as a function of applied voltage was recorded with 30 frames per second (fps) camera. The projected tip displacement (\ud835\udc62\ud835\udc43) of all the samples was later measured from recorded video using open source ImageJ software. The microstructure of CVD grown CNTs and various CNT/PDMS composite samples is studied through scanning electron microscopy (SEM) analysis in Fig. 7. SEM micrograph of as grown CNT shows their noodle like structure with an average diameter of 150 nm - 270 nm (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure7-1.png",
        "caption": "Fig. 7. (a) The active joint with first link (leg1) \u2013 conventional design; (b) the active joint with first link (leg 1) \u2013 modified design.",
        "texts": [
            " Links are assumed not to be rigid, so compliance for the links is obtained by materials science \u2013 structural analysis approach. A 3D spring model is assumed for the conventional and the modified design to incorporate all deflections along and around X, Y and Z axes of the Cartesian space. The compliance matrix is obtained by taking inverse of stiffness matrix in (3) (Section 3.2). LC = (SM) \u22121 (9) The compliance matrix in (9) represents the local stiffness matrix. It has to be transformed globally by rotational transformation as given below in (10). The method of rotation is shown in Fig. 7. R1 = \u23a1 \u23a2\u23a3 C\u03b81 \u2212S\u03b81 0 S\u03b81 C\u03b81 0 0 0 1 \u23a4 \u23a5\u23a6 \u00d7 \u23a1 \u23a2\u23a3 C\u03b41 0 S\u03b41 0 1 0 \u2212S\u03b41 0 C\u03b41 \u23a4 \u23a5\u23a6 . Rotation about Z axis Rotation about Y axis (10) For the conventional planar manipulator, \u03b41 is 0 (Fig. 7a). The following (11) determines the transformed link compliance. Equations (9) and (10) are substituted in (11) TCL11 = [ R1 0 0 R1 ] \u00d7 LC \u00d7 [ RT 1 0 0 RT 1 ] (11) The total stiffness CC12 is the combination of second joint stiffness and first link stiffness. Since the second joint is a passive joint, its compliance CJ12 is neglected. Thus, CC12 = TCL11 + CJ12 (12) Applying the previous method of transformation as in (8), TCC12 is obtained. The same procedure is repeated for the third mechanical module (third joint and its second link) CC12 = TCL11 + CJ12 (13) TCC13 = [ BRij BRij[ijPT ] 0 1 ] CC13 [ BRij 0 BRij[ijPT ] 1 ] (14) Adding (8), (12) and (14), the total compliance of R\u0304RR chain is given below"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002581_lra.2021.3070247-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002581_lra.2021.3070247-Figure2-1.png",
        "caption": "Fig. 2. Fabrication process of TCPOF. (a) Twisting of polymer-coated optical fiber. (b) Winding of twisted fiber onto mandrel. (c) Heating using an automatic oven. (d) Winding of heating wire onto fiber.",
        "texts": [
            " Section IV discusses the control of temperature and force by using the sensing function of the TCPOF under isometric conditions. Finally, Section V summarizes this study and discusses future research directions. Fig. 1 shows an overview of the TCPOF. As mentioned above, the TCPOF is fabricated using a polymer-coated optical fiber. In the TCPOF, the optical fiber performs the function of the sensor and the polymer coat mainly plays the role of the actuator. Length and temperature changes both affect the light intensity through the optical fiber, namely, the sensing function of the TCPOF, as discussed in Section III. Fig. 2 shows the fabrication process of the TCPOF. A nyloncoated optical fiber (GHN4001, Mitsubishi Chemical) was used as the TCPOF material because nylon is among the most frequently used materials for a TCPF and is relatively easy to obtain. Further, the optical fiber was fabricated using polymethyl methacrylate (PMMA). The diameters of the optical fiber and the coated fiber are 1.0 mm and 2.2 mm, respectively. Although the material cost is approximately five times that in the authors\u2019 previous study [16], this cost increase does not largely impair the cost of utilization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001369_aim43001.2020.9158850-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001369_aim43001.2020.9158850-Figure3-1.png",
        "caption": "Figure 3. Data augmentation process",
        "texts": [
            " To restrict the quaternion linear output (without using the activation function) to be unit quaternion, setting the hybrid loss is indispensable, including entropy loss for mode classification and quaternion loss for pose recognition, as below: Hybrid loss function Entropy loss for classification Quaternion loss Quaternion loss q q \ud835\udc64\u210e\ud835\udc52\ud835\udc5f\ud835\udc52 q \ud835\udc65 , \ud835\udc66 , \ud835\udc67 , \ud835\udc64 |\ud835\udc65 , \ud835\udc66 , \ud835\udc67 , \ud835\udc64 | 2 As in equation (2), shown above, the hybrid loss function tries to normalize the quaternion linear output and is used to approach the pose function set (ground true); the pose annotation system will be introduced to decide ground true quaternions. With respect to issue (2), ground true quaternions will be randomly multiplied by -1 to make the prediction more stable when the network is trained. Furthermore, it is highly recommended to conduct data augmentation while the network is being trained; that is, it should rotate the ground true pose and partial point cloud simultaneously with a random angle, as shown in Figure 3, or some noise should be added to the partial point cloud to stabilize the network. In this process, the z-axis of our robot\u2019s coordinate is set as the rotating axis, which is parallel to the desktop. Even though using quaternion as the pose representation presents several issues, it is still chosen as our representation because it has some beneficial attributes over other representations. For instance, quaternion can not only reduce the computational time when the network is trained but also avoid the gimbal lock problem, which may arise from using the Euler angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000503_012020-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000503_012020-Figure3-1.png",
        "caption": "Figure 3. Speed experienced by a water wheel when spinning",
        "texts": [
            " 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 Numerical analysis of the water wheel assumed that the waterwheel was half floating on the surface of the water, in accordance with the limitation of the problem that the incoming water flowing at a speed of 5 m/s then from the flow moves the windmill. The flow rate of the water that hit the blade on the water wheel causes the water wheel to rotate as shown in Figure 2 and Figure 3. In Figure 2 the red color is a waterwheel that gets pressure from the flow of water with a number of 12 pieces of blade. Where from the simulation results the pressure distribution experienced by each water wheel blade is evenly distributed, shown from the yellow color of each water wheel blade with an even pressure value of 43738.4 Pa. Figure 3 shows the results of the velocity distribution of water flowing on each blade of the waterwheel. The red color shows the highest flow velocity area on the waterwheel blade 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 with a maximum value of 3.42 m/s, while the lowest water flow velocity that affects the waterwheel blade is light blue with a value of 2.158 m/s. The results of numerical analysis show that from the simulation results of the waterwheel obtained at 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.1-1.png",
        "caption": "Fig. 57.1 Seat base loading and support points",
        "texts": [
            " The design requirement for a certain structural component is associated with certain attributes of the materials, such as, density, strength, and cost. Ashby [1] highlights the strategy of material selection into four main groups: (a) Translation of the design requirements, (b) screening of the materials using constraints, (c) ranking of the materials using the objective, and (d) seeking the supporting information. In the chosen problem, both strength and stiffness are important parameters for materials selection. Figure 57.1 presents the loads acting on the seat and support (load reaction) points of a seat base in a scooter. In order to simplify the approach to selectmaterials, the seat base can be conveniently considered as a flat panel subjected primarily to bending load (Refer Fig. 57.2), which can further be simplified as a beam bending in 1-D. Thus, the design requirements translate into the parameters as explained in Table 57.1. As thickness of the beam has been considered as the free variable in this case of loading, one could re-arrange Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000454_i-pact44901.2019.8960155-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000454_i-pact44901.2019.8960155-Figure11-1.png",
        "caption": "Fig. 11. Solid works model of the proposed robot",
        "texts": [
            " Nearly all types of DC motors have some internal mechanism, either electromechanical or electronic; to periodically change the direction of current flow in part of the motor. The Dc motor used is shown in the fig.10. In the proposed design, motors are used to rotate the wheels which provide locomotion to the robot. Four Dc motors are used for the movement of the wheels and two additional DC motors are used to provide movement to various parts of the gripper arrangement. With the help of solid works software the model of the proposed robot has been created. The solid work model is shown in fig. 11. V. RESULTS AND DISCUSSION A. Implementation The experimental robot shown in fig. 12 was developed, consisting of an autonomous robot facilitated with gripper arrangement. The output of ultrasonic sensor, soil moisture sensor and GPS module is shown in fig. 13, fig. 14 and fig. 15 respectively. Across the globe there are many ways that the robots could contribute, both economically as growing and harvesting more efficiently and cheaply, and ethically as increasing welfare via monitoring and timely intervention"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000922_pedstc49159.2020.9088409-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000922_pedstc49159.2020.9088409-Figure5-1.png",
        "caption": "Fig 5 : Phasor diagram of phase coils",
        "texts": [
            " The Ratio of rotor and stator teeth number is a determinant parameter in field flux path, in such a way that field flux can pass through the adjacent stator poles in 6/5 and 6/7 structure. This short flux path can lead to the better performance of machine like lower Total Harmonic Distortion (THD), in comparison with the long one. Windings and their connection are significant factor in VFRMs as individual phase coils should be connected in opposite directions when the number of rotor poles is odd. Fig. 5 shows the vectors of back-EMF in these ratios of rotor and stator teeth number. (a) (b) (a) (b) (c) (d) III. WINDING FUNCTION MODELING PROCEDURE OF A 6/7 VFRM In this section, the basic principal of WFM is going to be introduced. Turn, winding, and airgap function play key roles in winding function modeling of VFRMs. Turn function proposes the placement of windings among machine geometry. A 2-D cross-section view of a 6/7 VFRM is shown in Fig. 6. The windings of Phase a, phase b, and phase c, are in red, green, and blue, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000149_1350650119887043-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000149_1350650119887043-Figure4-1.png",
        "caption": "Figure 4. Experimental set-up for foil bearing clearance measurement.",
        "texts": [
            " On one end of the shaft, air turbines drive is mounted to spin the shaft at desired speed by passing the compressed air over the turbines. On the other end of the shaft, a sleeve of 10mm inner diameter is introduced whose outer diameter can be varied depending on the test bearing inner diameter. This makes the test rig versatile to accommodate bearings of different sizes with minimal modification. The sleeve used here has a tolerance of 10 mm and ground surface finish. Experimental set-up for the clearance measurement is shown in Figure 4. Provision is made to load the bearing in both vertically upward and downward direction for clearance measurement. Eddy current displacement probe is mounted vertically on a firmly mounted flat plate for gap measurement. Initially, the bearing is loaded in the downward direction up to 1 kg load in steps of 500 g and the readings shown by the displacement probe are noted. Under the loaded condition, the bearing is exerted with 2 kg load in steps of 500 g in reverse direction resulting in a net load of 1 kg in an upward direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001706_s12555-020-0031-7-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001706_s12555-020-0031-7-Figure2-1.png",
        "caption": "Fig. 2. Connected links exert forces and moments through the joints.",
        "texts": [
            " The issues pertaining to the discrete implementation of the proposed control scheme is discussed in Section 6. The results of simulation experiments are presented in Section 7 along with a detailed discussion, and finally, the paper is concluded in Section 8. A serial-link manipulator consists of several links with joints that allow relative motion between them. Apart from allowing motion, links also physically interact in terms of interactive forces and moments between them through these joints. As illustrated in Fig. 2, the link (i\u2212 1) exerts a force fi and a moment ni on the link i. Similarly, the link (i+ 1) exerts a force \u2212 fi+1 and a moment \u2212ni+1 on the link i. With these interactive forces and moments from connected links, the ith link experiences a net force of Fi and a net moment of Ni. The actuator applies a moment (or force in the case of prismatic joints) about (along) the joint axis Zi. We may consider a \u2018joint-link pair\u2019 as a subsystem or an agent, interacting with other subsystems/agents. In this sense, a serial-link robot is a multi-agent system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003241_ldia49489.2021.9505999-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003241_ldia49489.2021.9505999-Figure1-1.png",
        "caption": "Fig. 1. Comparison of two kinds of WWCEC systems",
        "texts": [
            " Independent wind or wave power generation systems that can convert wind energy or wave energy separately have been investigated [3, 4], and they are going to be put into production. However, too many independent systems result in a huge of sea area occupations and high development cost. Additionally, it is difficult for independent energy conversion systems to continuously and stably output electric energy [5, 6]. Hence wind-wave combined energy conversion (WWCEC) systems have arouse a lot of interests recently [5, 7-9]. In WWCEC system, both energies can be converted on one platform so that the system development cost and the nonworking hours can be reduced. Fig. 1(a) presents the diagram of the platform-shared WWCEC system [7, 8] which is one type of WWCEC system. However, multiple generators are needed to work at the same time in this system so that the system volume is huge. The other common WWCEC system is mechanical coupling type WWCEC system [9] which needs too many mechanical devices so that this system generally has low efficiency, poor stability and high maintenance cost. It can be seen that traditional WWCEC systems have several disadvantages, which restrict the large-scale application of WWCEC system. From this scholars creatively proposed a new direct-drive WWCEC system based on two-degree-offreedom linear-rotary generators (LRGs) [10-14], which is shown in Fig. 1(b). The collected kinetic energy can be directly converted to electric energy by the LRG, so the novel WWCEC system has the advantages of high efficiency, high reliability, good stability and low cost. It is going to be an alternative candidate for offshore wind energy and wave energy conversion systems. Thanks to the application of LRG, the novel WWCEC system does not require multiple generators to work, and avoids too many mechanical devices. Many scholars begin to research on LRGs. In [14], L"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000264_icems.2019.8921522-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000264_icems.2019.8921522-Figure7-1.png",
        "caption": "Fig. 7. Flowchart of vibration analysis.",
        "texts": [
            " The vibration analysis of the stator of the PMaSynRel is divided into two parts, which are the modal analysis and harmonic response analysis. Firstly, the modal analysis is aiming to compute the dynamic characteristics (natural vibration modes or breathing modes) of the stator. And then, the harmonic response analysis will be conducted by introducing the electromagnetic forces from Maxwell 2D simulation. In the vibration analysis, the electromagnetic forces works as vibration excitation [17]. The flowchart of the vibration analysis has been summarized, as shown in Fig. 7. A. Modal Analysis Modal analysis is done in ANSYS Workbench to get the structural characteristics of the PMaSynRel motor, in terms of the natural frequency, modal modes and the mode shape. The results of the modal analysis of the stator and the rotor are listed in Table II and Table III, respectively. According to the modal shapes of the circumferential modes shown in Table II and Table III, there is a great influence on both x- and y- axes caused by the potential vibration at the corresponding natural frequencies",
            " Furthermore, the modal analysis provides a good basement for the future analysis, such as the harmonic response analysis. In addition, it is useful for the designer to get an idea about the harmonic orders of the electromagnetic excitations which excites any of those vibration modes and result peaks of the NVH response, which will be discussed in detail in the next subsection B. Hence, the designer can define which harmonic orders needed to be mitigated in the electromagnetic design process of the motor. As presented in Fig. 7, the multi-physic model of the PMaSynRel motor is established and calculated based on a 2D electromagnetic model in Maxwell. At the beginning, the electromagnetic force, which is acting on the tooth tips of the stator, is calculated and obtained in Maxwell. Then, the Maxwell 2D project is connected to the harmonic response analysis project in Ansys Workbench, so the electromagnetic force is imported to apply on the surface of each tooth tip of the stator. Finally, the vibration simulation is run by using superposition method, which means that the modal analyses simulations have been applied firstly, then, the time harmonic response simulation has been conducted after"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001154_rpj-09-2018-0243-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001154_rpj-09-2018-0243-Figure3-1.png",
        "caption": "Figure 3 (a) Schematic of sample preparation for measuring surface roughness (i.e. shown for the (Z) build oriented sample); (b) Experimental setup for surface roughness measurements of DMLS IN718 specimens using Veeco Dektak3ST surface roughness profilometer",
        "texts": [
            " To assess the relationship between surface roughness, build orientation and torsional fatigue fracture response behavior of as-built DMLS IN718, average and root mean square surface roughness measurements were taken using the Veeco Dektak3ST contact surface profilometer. A slice from the outer gauge section of the all specimens (X, Y, Z, XZ45, YZ45 and XY45) was taken using the low speed Isomet cutter and used to determine the surface profile. The sectioning process is as depicted in the schematic in Figure 3a. Three measurements were taken at room temperature on the as-built section of each sample using a 5 mm diamond stylus tip, over a scan length of 500 mm, incorporating 1,000 data points, for a horizontal scan resolution of 0.5 mm/sample. The scan profile mode was set to \u201cHill,\u201d and the scan speed was kept at medium settings. The experimental setup of the specimens during surface roughness measurements is as shown in Figure 3b. Measurements were taken toward the center of the as-built surface to reduce edge effects. Manufactured Inconel 718 Sanna F. Siddiqui and Ali P. Gordon Rapid Prototyping Journal Manufactured Inconel 718 Sanna F. Siddiqui and Ali P. Gordon Rapid Prototyping Journal DMLS IN718 specimens manufactured along each build orientation were subject to room temperature torsional fatigue tests, under the same experimental conditions as its conventionally manufactured counterpart. Completely reversed (Rf = 1) torsional fatigue tests were performed at an angle of twist cycling range of Df = 615\u00b0 and twisting rate of 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure3-1.png",
        "caption": "Figure 3. Design of the bolted (a) and rivet bolt (b) connection of the MC-21 aircraft compartment",
        "texts": [
            " 1) and when connecting the panels along the orbital joints in the process of assembling the fuselage (Fig. 2), bolt-ribbing and, less often, bolted joints. ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 The design of the stud bolt and bolt connections varies depending on the type of rod and bolt head, the size of the head, the material used for the fixing element, and the method of installation in the hole. Examples of rivet and bolt connections are shown in Figure 3 [7]. ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 As a rule, an aircraft's overall resource is determined by the definition of its frame resource. In case of damage to the frame, failures occur that make up 12...30% of the total number of failures [3, 4]. The main cause of failures in most cases is fatigue damage, which occurs at the joints of parts and assemblies [3, 5, 6]. Connection weariness is determined by the following factors [8]:  peculiarity of the aircraft design and rationality of power configuration (number of rows of fasteners and distribution of loads between them);  the value of stresses in the elements of the junction;  the peculiarity of the assembly process (in particular, the axial and radial interference in the connection)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001044_s12206-020-0533-5-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001044_s12206-020-0533-5-Figure11-1.png",
        "caption": "Fig. 11. The graded circular plate.",
        "texts": [
            " 8 and 9, the robot exhibited a free periodic motion at a period of 1.5 s between start position A and end position B. In addition, for the time interval of 0-0.75 s, the re- sults of the Simmechanics model are well in accordance with the results of the TPBVP (Figs. 4 and 5). A TLRM with adjustable counterweights manufactured in Semnan Robotics Lab was used for the experimental implementation. The robot construction with associated counterweights are demonstrated in Fig. 10. A graded plate was used to adjust the counterweight arm angle (Fig. 11). Components of the counterweight are presented in Fig. 12. The servo- actuators used in the robot were Dynamixle XH430-W210 by Robotis\u00a9. Indeed, due to friction, air resistance, and parametric uncertainties, it was practically difficult to implement the motion repeatedly without consuming any power as an open-loop policy to exactly reproduce the situation in the previous section. Thus, in order to implement the proposed method, the servoactuators were operated in position-control mode and the experimental apparatus was set up following a closed-loop approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure3-1.png",
        "caption": "Fig. 3. Hertz contact between ball and raceway: (a) contact between steel ball and raceway; (b) contact parameters a, b, \u03b4.",
        "texts": [
            " The discussion in this article can be used as a supplement to the deformation and displacement derivation process of the static analysis model, and provide a reference for bearing users to calculate the nonlinear deformation of the bearing. Hertz\u2019s theory is widely used to analyze the contact between ball and raceway. According to the Hertz contact theory, the contact surface where the steel ball contacts the raceway is an ellipse, and its load distribution is an ellipsoid in space; the maximum contact pressure and deformation occurs at the center of the contact ellipse, as shown in Fig. 3. The normal contact deformation \u03b4 is determined by Eq. (1) [7]: 2 3 * 3 2 2 Q E \u03c1 \u03b4 \u03b4 \u03c1 \u239b \u239e = \u239c \u239f\u239c \u239f \u239d \u23a0\u2032 \u2211 \u2211 (1) (for the meaning of each parameter, see the Nomenclature). For ball bearings, the steel ball usually needs to be in contact with the inner raceway and the outer raceway at the same time to transmit the load. So, the total normal deformation is the sum of the inner and outer normal contact deformations. See Eq. (2). n i e\u03b4 \u03b4 \u03b4= + 2 2 3 3 * *3 3 2 2 2 2 i e i e i e Q Q E E \u03c1 \u03c1 \u03b4 \u03b4 \u03c1 \u03c1 \u239b \u239e \u239b \u239e = +\u239c \u239f \u239c \u239f\u239c \u239f \u239c \u239f \u239d \u23a0 \u239d \u23a0\u2032 \u2032 \u2211 \u2211 \u2211 \u2211 (2) 2 2 3 3 2 * * 33 3 2 2 2 2 i e i e i e Q E E \u03c1 \u03c1 \u03b4 \u03b4 \u03c1 \u03c1 \u23a1 \u23a4 \u239b \u239e \u239b \u239e\u23a2 \u23a5= +\u239c \u239f \u239c \u239f\u23a2 \u23a5\u239c \u239f \u239c \u239f \u239d \u23a0 \u239d \u23a0\u23a2 \u23a5\u23a3 \u23a6 \u2032 \u2032 \u2211 \u2211 \u2211 \u2211 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003191_itec51675.2021.9490116-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003191_itec51675.2021.9490116-Figure1-1.png",
        "caption": "Fig. 1. Cross-section of the proposed five-phase synchronous reluctance motor.",
        "texts": [
            " Finally, the overall electromagnetic performance and effect of the third harmonic current component will be verified by finite element analysis. It will be shown that the torque density can be improved by utilizing the harmonic current injection scheme and the proposed five-phase synchronous reluctance motor is an appealing non-PM candidate for taction applications. The machine topology and drive system configuration of the proposed five-phase synchronous reluctance motor drive is introduced in this section. Fig. 1 shows cross-section, power converter, and drives topologies for the proposed fivephase synchronous reluctance motor. The five-phase symmetric distributed windings and simple salient reluctance rotor are employed in the proposed synchronous reluctance motor. The rotor structure selection will be discussed in the next section, which will show that the simple salient reluctance rotor is more suitable for the proposed five-phase synchronous reluctance motor because of its good coupling capability between fundamental and third harmonic MMF [51]-[52]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000530_iros40897.2019.8967655-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000530_iros40897.2019.8967655-Figure1-1.png",
        "caption": "Fig. 1: Geometrical design of selected sensor. The 3D crosssectional view shows one stretched EGaIn microchannel.",
        "texts": [
            "au Of particular interest in this paper are EGaIn-based strain sensors which have been designed in several types and shapes for both uniaxial [14] and multiaxial [15] measurements.These sensors have also been designed to be embedded in the SPAs [16], or mounted externally or as a sensory skin [17]\u2013[20]. Our study focuses on using uniaxial EGaIn strain sensors in estimating the strain based on the measured resistance. After reviewing the complexity and manufacturing cost of several EGaIn-based soft sensors in the literature, one design was selected for our study as shown in Figure 1 and was fabricated according to the manufacturing technique presented in [17]. The contributions of this paper are to 1) introduce a new mathematical model to estimate the strain as a function of resistance and determine the estimation accuracy, 2) study the effect of strain rate on modelling accuracy using a custom procedure designed for analysing hysteresis in soft sensors. The reminder of the paper is organised as follows. Section II presents related work, followed by a description of sensor geometry in Section III",
            " This accuracy is required if the proprioceptive sensing is expected to replace the external perception when detecting SPAs deformation. However, less accuracy might be required depending on the targeted applications [17]. We extends the existing literature by introducing a new model, and focus on studying its accuracy limits, and the effect of strain rate (or speed) on hysteresis. This could lead the way to understand the possibility of reconstructing the deformation of soft actuators using a network of sensors in future work. The selected sensor, shown in Figure 1, has an active length of Lcc = 76.3mm between the connection centre points where the strain will be applied while the remaining length is inactive. The active length has two main regions. The first region consists of two distinctive subregions in which one represents the sensing element with a length of Ls = 30.0mm and the other contains connection microchannels with a length of Le = 18.0mm. This enables the connection of measuring devices to read sensor resistance. The second region represents the remaining length within Lcc that has no microchannels in it, and transfers applied strain from centre points to the sensing element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002736_s13177-021-00256-3-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002736_s13177-021-00256-3-Figure2-1.png",
        "caption": "Fig. 2 Kinematic model of non-holonomic vehicle",
        "texts": [
            " 1 A(xA, yA), J(xJ, yJ), B(xB, yB), D(xD, yD) and E(xE, yE) points are fixed to the vehicle where A, J, B, D points are the corners and E point is the center of the rear track of the vehicle. Additionally, G(xG, yG) and F(xF, yF) points are the lower-left and the upper-right corner of the parking spot. (See Fig. 1). Also, all the methods in this paper are presented for the scenario where the parking spot is on the right side. In order to park the vehicle to its left, same equations can be modified using symmetry. Since the parking maneuver is performed at low speeds, kinematic equations provide the motion model sufficiently accurate. Figure 2 shows the kinematic model where (x, y) denote the position of center of the rear axle and \u03c8, \u03b4 stands for the orientation angle and the steering angle of the vehicle respectively. The standard kinematic equations of the vehicle are given in Eq. (1) [18] where a denotes the wheelbase of the vehicle and u1, u2 stands for the linear velocity and the steering velocity of the vehicle respectively. x\u0307 \u00bc u1cos \u03c8\u00f0 \u00de y\u0307 \u00bc u1sin \u03c8\u00f0 \u00de \u03c8\u0307 \u00bc u1 a tan \u03b4\u00f0 \u00de \u03b4\u0307 \u00bc u2 8>>< >>: \u00f01\u00de According to the Ackermann, instantaneous center of rotation (ICR) for left and right steering are represented as Cl and Cr as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000126_icca.2019.8900006-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000126_icca.2019.8900006-Figure1-1.png",
        "caption": "Fig. 1. An overview of the suspension system.",
        "texts": [
            " In this section, we describe the suspension system of medium-speed maglev trains and give an approximate dynamic model, then discuss about it. The electromagnetic suspension system of the maglev train depends on the attraction between rail tracks and the electromagnet mounted on the vehicle body, which can be considered as the components of magnetic suspension system. The suspension system of the entire train consists of multiple electromagnets, but the decoupling can be realized between multiple electromagnets. Therefore, the multi-point control system is simplified to a single point control. Fig. 1 shows an overview of the suspension system and its detail. For the convenience of theoretical analysis, a simplified version is given in Fig. 2. The simplified suspension system includes two main parts, a single electromagnet and the rail track. F (i(t), x(t)) indicates the electromagnetic force between the electromagnet and the track; A denotes the effective magnetic pole area of the electromagnet; N denotes the number of turns of the coil; i(t) and u(t) denotes respectively the current and voltage in the coil of an electromagnet; f denotes the disturbance mainly contributed by the mass of passengers; R denotes the coil resistance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001429_j.procir.2020.05.180-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001429_j.procir.2020.05.180-Figure1-1.png",
        "caption": "Fig. 1. A part with Workpiece Coordinate System (WCS) and a Feature Coordinate System (FCS) for a cylinder.",
        "texts": [
            " Regardless of the coordinate system, the position of any point in space can be represented by a position vector OP which defines the location of the point P with reference to the origin. In manufacturing applications, separate coordinate systems may be defined for each manufacturing feature or functional surface to facilitate local process planning such as machining operations. The feature coordinate system (FCS) is defined with respect to the workpiece coordinate system (WCS) and may be oriented differently as displayed in Fig. 1. The location of a directional vector in 3 is typically defined by a translation vector 3t , while the orientation may be represented as a 3 3 special orthogonal matrix ( (3))SOR . It is common to combine the rotation matrix R and the translation vector t in a single homogeneous transformation matrix (3)SET : 1 30 1 0 0 0 1 x x y z y z t r r r t t                 R t T (1) The representation of a vector in 3 may thus be condensed into a representation in the form of ( , , , , , )x y z x y zr r r t t t "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003188_s13632-021-00748-4-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003188_s13632-021-00748-4-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the holder used for pressurizing the TLP bonding assemblies in the furnace.",
        "texts": [
            " For this purpose, different TLP bonds were fabricated using BNi-2 interlayer. In this study, Inconel 617 and stainless steel 310 were machined out of the as-received materials to 40\u00d710\u00d75 mm dimensions and BNi-2 interlayer with thickness of 40 \u00b5m and with the same dimension of the base metals was used for joining the base metals. Chemical compositions of the base metals and the interlayer are listed in Table\u00a01. The steel sample holder used for holding and pressurizing the TLP bonding assemblies in programmable resistance furnace is schematically shown in Fig.\u00a01. The specimens were first polished down to 1200 grit sand paper and were then ultrasonically cleaned with acetone for 30 min with a frequency of 35 kHz at 0.5 A before being carefully fixed in the holder. TLP bonding process was carried out at different temperatures of 1040, 1070 and 1100 \u00b0C for 60 min under a vacuum of 105 mbar. Samples were cooled inside the furnace to ambient temperature after the bonding process. Heating and cooling rate of the furnace before and after the actual TLP process was set to 10 \u00b0C/min"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002548_s12555-020-0744-7-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002548_s12555-020-0744-7-Figure1-1.png",
        "caption": "Fig. 1. Sketch of a torque\u2013driven inertia wheel pendulum.",
        "texts": [
            " Simulations results on an inertia wheel pendulum model are given in Section 4. Finally, we offer some concluding remarks in Section 5. 2. A TORQUE\u2013DRIVEN INERTIA WHEEL PENDULUM The torque\u2013driven inertia wheel pendulum is an underactuated mechanical system consisting of a physical frictionless rigid pendulum rod with a symmetric disk (wheel) attached to the distal tip, which is free to spin (thanks to an ideal torque actuator) about an axis parallel to the axis of rotation of the pendulum rod (see Fig. 1). 2.1. A plant model The energy function (Hamiltonian) is the sum of the kinetic energy plus potential energy of the mechanical system. With regard to the inertia wheel pendulum plant, it is given by [12] H(q, p)= 1 2 [a1 p2 1+2a2 p1 p2+a3 p2 2]+m3[cos(q1)\u22121], (1) where q = [q1 q2] T = [\u03b81 \u03b82] T and p = [p1 p2] T are the vectors of generalized positions and momenta, respectively, the M inertia matrix is given by M = [ a1 a2 a2 a3 ] , (2) and the potential energy function U(q1,q2) = m3[cos(q1)\u22121] (3) being a1 = I1 + I2, a2 = I2, a3 = I2, m3 4 = g(m1lc1 +m2l), where we have assumed (I1 + I2) >> (m1l2 c1 + m2l2) in order to simplify the model. From Fig. 1, \u03b81 and \u03b82 are the joint positions of the pendulum and the wheel, respectively, and u is the control torque input acting between wheel and pendulum rod. We has been assumed both mechanism joints without friction. Table 1 shows the meaning of the remaining parameters. In Hamiltonian formulation the momentum p is defined as [29] p = Mq\u0307, (4) where q\u0307 is the vector of generalized velocities. The dynamic model of the torque\u2013driven inertia wheel pendulum without viscous friction, can be written as d dt [ q p ] = [ 02\u00d72 I2\u00d72 \u2212I2\u00d72 02\u00d72 ][ \u2207qH(q, p) \u2207pH(q, p) ] + [ 02 G ] u, (5) where G = [ 0 1 ] , (6) and u is the torque control input, which is assumed to be provided by an ideal perfect memoryless actuator",
            " Next, since T\u22121 = MK\u22121M\u22121 a in accordance with (14), (31) can only be expressed in terms of the qa variable G\u22a5MK\u22121M\u22121 a \u2207qa Ua(qa) = G\u22a5\u2207qU(q), (41) once the right-hand side of (41) is first computed and then evaluated for qa and qd(t), that is G\u22a5\u2207qU(q)|(q=K\u22121qa+qd(t)) , (42) where q is written in accordance with (10), and qd(t) = [qd1(t) qd2(t)] T , such that a solution Ua(qa) may be obtained from (41). Therefore, consider (10)-(11) with K and Ma given by (16), and the G\u22a5 = [ 1 0 ] matrix defined in Proposition 1, then the PDE in (41) yields \u03b11 \u2202Ua(qa) \u2202qa1 +\u03b12 \u2202Ua(qa) \u2202qa2 =\u2212m3det[Ma]sin ( qa1 a1 + q\u0304d1 ) , (43) where \u03b11 = [d3\u2212 d2], \u03b12 = [d1\u2212 d2], and the right-hand side of (43) was computed in accordance with (42), being qd1(t) = q\u0304d1 with the constant q\u0304d1 \u2208 {. . . ,\u22124\u03c0,\u22122\u03c0,0,2\u03c0,4\u03c0, . . .} (44) corresponding to the upright desired positions of the pendulum (see Fig. 1). Because (43) is an easy linear PDE, a solution for (43) is given by Ua(qa) = a1m3det[Ma] [d3\u2212d2] [ cos ( qa1 a1 + q\u0304d1 ) \u22121 ] + kp 2 [qa2 \u2212 \u03b32qa1 ] 2, (45) where \u03b32 = (d1\u2212d2) (d3\u2212d2) , and to guarantee positivity of Ua, the di elements are chosen to hold the inequalities d1 > d3, and d2 > d3, (46) such that d1d3 \u2212 d2 2 > 0, and kp > 0. Moreover, Ua is positive definite with respect to the variables given by [(qa1 \u2212a1[\u03b4\u03c0\u2212 q\u0304d1 ]) (qa2 \u2212 \u03b32a1[\u03b4\u03c0\u2212 q\u0304d1 ])] T for any \u03b4 even number. Once the Ma and Ua solutions are obtained, we write the control law (32) following the notation of [12] as follows: u = ues +udi, (47) where ues is the potential energy shaping term given by ues = [GT G]\u22121GT [ \u2207qU(q)\u2212MK\u22121M\u22121 a \u2207qa Ua(qa) ] , = 1 det[Ma] [[ \u2212a2 a1 d3 + a3 a2 d2 ] \u2202Ua(qa) \u2202qa1 + [ a2 a1 d2\u2212 a3 a2 d1 ] \u2202Ua(qa) \u2202qa2 ] , (48) with GT \u2207qU(q1) = 0 in accordance with (2)-(3), and \u2202Ua(qa) \u2202qa1 =\u2212 m3det[Ma] [d3\u2212d2] sin ( qa1 a1 + q\u0304d1 ) \u2212 \u03b32kp[qa2 \u2212 \u03b32qa1 ], \u2202Ua(qa) \u2202qa2 =kp[qa2 \u2212 \u03b32qa1 ]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure17-1.png",
        "caption": "Fig. 17. (a) Generated Mesh for the commutator gap. (b) Generated Mesh for the lateral gap on the window side. (c) Magnified view of a portion of mesh. Boundary conditions imposed on commutator and lateral gaps are also presented.",
        "texts": [
            " Utilizing the mesh capabilities of OpenFOAM such as mesh rotation and snappyHexMesh [31], a mesh is generated at different angular steps upon which the Reynolds Equation is solved for pressure distribution. The nodes and elements (cell centers) of the mesh are rotated by discrete angles using OpenFOAM capabilities. SnappyHexMesh utilizes the Stereo Lithography (.stl) files of ports to remove the port regions, where the flow is turbulent, from the background mesh. Sample meshes of commutator gap and lateral gap are presented in Fig. 17. Mesh sensitivity analysis is performed to confirm that the results are not mesh sensitive beyond 2.5 million cells in the lateral gap and 0.5 million cells in the commutator gap. Dynamic pressure values extracted from the lumped parameter fluid dynamic model are provided as boundary conditions to the gap model. The boundaries of mesh where the suction, delivery and chamber pressure boundary conditions are applied are represented in Fig. 17. The exterior of the mesh which corresponds to the polymer seal is subjected to a solid wall boundary condition which implies a zero-pressure gradient. Leakages are evaluated as a derived quantity of the pressure distribution In the proposed approach, the deformation of solid components due to contact is considered because it has a direct impact on the torque. Whereas, the deformation of components because of pressure within the lubricating gaps has a secondary effect on the torque. Since the constant undeformed gap height assumption provided good agreement of volumetric efficiency with experimental values in the past work [12], the same assumption is considered in this study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002248_j.matpr.2020.07.608-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002248_j.matpr.2020.07.608-Figure2-1.png",
        "caption": "Fig. 2. 3R Planar Manipulator [14]",
        "texts": [
            " In the above algorithm in step 5, while updating joint angle values, the inverse of the jacobian matrix is required to be calculated. The inverse can be defined only for an invertible matrix, i.e., the square and non-singular matrix. The extension of inverse operation for non-singular matrices is known as the Moore-Penrose pseudo-inverse or pseudo-inverse. If A is invertible, then the Moore-Penrose pseudo-inverse is equal to the matrix inverse. The inscribed figure of developed manipulator is shown in Fig. 2 and parameter Table 1 shown: Mathematical formula: Fig. 1. Diagram of the Newton-Raphson Algorithm used for IK Calculations. JTJDh \u00bc JT e! After solving inverse kinematics to find required joint angles, we can write Dh \u00bc J\u00fe e! J\u00fe \u00bc JT\u00f0JJT\u00de 1 0T1 \u00bc cosh1 sinh10l1cosh1 sinh1cosh10l1sinh1 0010 0001 2 6664 3 7775 \u00f0a\u00de 1 T2 \u00bc cosh2 sinh20l2cosh2 sinh2cosh20l2sinh2 0010 0001 2 6664 3 7775 \u00f0b\u00de 2 T3 \u00bc cosh3 sinh30l3cosh3 sinh3cosh30l3sinh3 0010 0001 2 6664 3 7775 \u00f0c\u00de Thus, 0T3 = 0 T1 1 T2 2 T3 The equation gives a transformation matrix of the end-effector\u2019s frame w",
            " l1 \u00bc 2units l2 \u00bc 2units l3 \u00bc 2units The workspace of the manipulator is a circular plane of radius 6 units. Considering origin on the base frame of manipulator, inverse kinematics can be solved for each point in the workspace. Consider a point (2,4) within the workspace and if we solve for corresponding joint angles of manipulator links, the algorithm converges iteratively and finally, we get solution as: First joint angleh1 = 88.41 Second joint angle h2 = 88.36 Third joint angleh3 = 91.61 All angles are in degree.Fig. 2. The pseudo-inverse method performs well inside the workspace. It converges to the given target point (2, 4) iteratively as shown in Fig. 3: This method performs poorly near singular configurations due to instability, as shown in Fig. 4: Damped least-squares using Singular value decomposition For the investigation of damped least square methods, Singular Value Decomposition (SVD) comes out as a powerful technique. A singular value decomposition of J comprises of stating J as the representation VT Where U and V are orthogonal matrices and D is a diagonal matrix with diagonal entries ri: The singular value decomposition of J mainly subsists and entails that J can be represented as Xr i\u00bc0 UDVT Hence on simplification, we get the final damped least square solution in the form Dh \u00bc \u00f0 Xr i\u00bc0 ri ri 2 \u00fe k2 viui T\u00de e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001026_ab9549-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001026_ab9549-Figure14-1.png",
        "caption": "Fig. 14. (a) PVC frame and pre-stretched PDMS sheet, (b) PVC frame bonded onto pre-stretched PDMS sheet, (c) formation of curved beam structure at new equilibrium position after removal of pre-stretching force, (d) conductive grease electrodes applied on both sides of elastomer. (e) Actuation behaviour on the application of voltage, (f) dimension of frame and pre-stretched elastomer. S is the length of rectangular cavity and b/2 is the width of the frame.",
        "texts": [
            " \ud835\udf02 is defined as the ratio of mechanical energy stored (\ud835\udc48A) to the electrical energy supplied to the actuator (\ud835\udc48c), \ud835\udf02(%) = \ud835\udc48\ud835\udc34 \ud835\udc48c \u00d7 100. (1) The electrical energy supplied in a capacitive DEA is, \ud835\udc48c = 1 2 \ud835\udc36\ud835\udc492, (2) where C = \ud835\udf160\ud835\udf16(A/d) with A and d are the surface area and thickness of the elastomer film. Knowing the geometric parameter of DEA and \ud835\udf16 from Fig. 9(a), it is straight forward to calculate \ud835\udc48\ud835\udc50 at an applied voltage V from eq. (2). However, the calculation of UA is not that obvious. During the fabrication of DEA, a pre-stretched elastomer film is bonded to PVC sheet (Fig. 14(a), (b)). DEA takes the shape of a curved beam structure after removing the pre-stretching force [73] as shown in Fig. 14(c-d). On applying voltage, compressive Maxwell stress develops along the thickness of elastomer layer. Poisson\u2019s effect causes the lateral expansion in elastomer and hence actuation of DEAs (Fig. 14(e)). To calculate UA, it is assumed that DEA is deforming under constant bending moment due to the fact that at any instant, the actuator remains in a curved shape having fix radius of curvature (Fig. 15). DEA consists of a PVC frame with a rectangular cavity of length S and flange width b/2. The width of actuator is w. The thickness of the frame and elastomer is d1 and d2 respectively (Fig. 14(f)). Let the initial and final radius of curvature of actuator are R0 and R. The bending moment M of curved beam is given by the relation, \ud835\udc40 = \ud835\udc38\ud835\udc3c ( 1 \ud835\udc45 \u2212 1 \ud835\udc450 ) , (3) and the associated bending energy U is, \ud835\udc48 = \ud835\udc402 2\ud835\udc38\ud835\udc3c \u222b \ud835\udc51\ud835\udc46 \ud835\udc46 0 = \ud835\udc402 2\ud835\udc38\ud835\udc3c \ud835\udc46. (4) Combining eqs. (3) and (4) provides U as a function of radius of curvature, \ud835\udc48 = \ud835\udc38\ud835\udc3c 2 ( 1 \ud835\udc45 \u2212 1 \ud835\udc450 ) 2 \ud835\udc46. (5) 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d M nu sc rip t 11 Using eq",
            " \ud835\udc4a = \ud835\udf070 2 (\ud835\udc3c1\u0305 \u2212 3) + \ud835\udc3e0 2 (\ud835\udc3d \u2212 1)2, (7) where \u00b50 is shear modulus and K0 is bulk modulus. These material parameters are calculated from elastic modulus and Poisson\u2019s ratio obtained in section 3.2. \ud835\udc3c1\u0305 is the first deviatoric strain invariant \ud835\udc3c1\u0305 = \ud835\udc3d\u2212 2 3(\ud835\udf061 2 + \ud835\udf062 2 + \ud835\udf063 2) and \ud835\udc3d = \ud835\udf061\ud835\udf062\ud835\udf063. Here, \ud835\udf061, \ud835\udf062, \ud835\udf063 are the stretches in elastomer film along three directions. In line with the experiments, 30% prestretch is applied to elastomer film before bonding it to PVC frame and the assembly is allowed to relax. This leads to curved beam structure of DEA as shown in Fig. 14 (e) and Fig. 15(a). The DEA stress \ud835\udf0e\ud835\udc60 as a function of applied electric field Ef is given by the relation [32], \ud835\udf0e\ud835\udc60 = \ud835\udf061 \ud835\udf15\ud835\udc4a\ud835\udc60\ud835\udc61\ud835\udc5f\ud835\udc52\ud835\udc61\ud835\udc50\u210e \ud835\udf15\ud835\udf061 \u2212 \ud835\udf160\ud835\udf16\ud835\udc38\ud835\udc53 2. (8) \ud835\udc4astretch is the strain energy density W associated with prestretching of elastomer film and \ud835\udf160\ud835\udf16\ud835\udc38\ud835\udc53 2 is induced Maxwell stress. A compressive Maxwell stress ( \u2212 1 2 \ud835\udf160\ud835\udf16\ud835\udc38\ud835\udc53 2) is applied to elastomer film to simulate the actuation behaviour of DEA. The deformed shapes of DEA under applied stress are shown in Fig. 16 at different simulation time steps. The energy stored in DEA obtained from FE simulations is compared with UA calculated using eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001102_j.matpr.2020.05.310-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001102_j.matpr.2020.05.310-Figure3-1.png",
        "caption": "Fig. 3. Load application on the wheel housing.",
        "texts": [
            " For structural mechanics problems it is observed that the finite element characteristics are identical whichever of the two methods is used. The wheel housing designed was meshed in the software for proceeding with the nodal analysis to be performed in the ANSYS software [23]. The meshed view is shown in the Fig. 2. The mesh is then opened in the software for the analysis in which initially the load is applied to check for the stresses. The load applied gets enforced in the housing which is portrayed in the Fig. 3. On applying the stresses in the housing the deflection is noted in the output as 0.42 mm for the applied load as shown in the Fig. 4. The stress acting on the housing of the wheel while applying load is 39.008 N/mm2 as al., FEA based approach on replacing the metal cast wheel into thermoset 20.05.310 shown in the Fig. 5. The values determined were found to be within the safety limits since the maximum permissible value was 118 N/ mm2 mentioning consideration in usage of the made wheel housing. As an outset, considering various calculations and design development for the product and mould design for SMC, wheel housing was completed with actual values as guidance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure3-1.png",
        "caption": "Fig. 3. Comparison of Mesh quality plots at room temperature.",
        "texts": [],
        "surrounding_texts": [
            "The temperature distribution on the base alloy at 293 K is shown in Fig. 2. While comparison of mesh quality plots is indicated in Figs. 3 and 4. This was done at room temperature in order to pre-heat the base metal and check the stresses via modelling. Pre-heat treatment of base alloy is good because it allows it to absorb the laser beam uniformly and avoid residual stresses that can tamper with the microstructures with larger grain sizes and influence the surface quality of the additive manufactured component. In addendum, component load resistance can be tampered with when the residual stresses are larger. Standard quality requirements must be met, and permissible irregularities must be identified and categorized. Different kind of anomalies and their influence on the additive manufactured surface performance and quality must be well understood. The wrong combinations of power, thickness of layer, direction of build, scan velocity and rate of flow can impart on the anomalies. The chief challenges debarring the full implementation of additive manufacturing is quality assurance. The real quality assurance involves predictive modelling and real time techniques of detecting anomalies immediately in additive manufacturing. Thermal cycle in additive manufacturing of metals impart residual stresses in additive manufactured parts and this is a real concern that can lead to parts distortion. The flow of molten liquid in the melt pool that came about as a result of the increase in the speed of the coating materials was linked to higher laser power. Standard distribution, enhancement and spheroidization are linked to higher laser power as shown in Figs. 5\u20137. Movement of laser from the starting point to the end changes the temperature distribution and enhanced microstructures in the melt pool are depended on the laser input, scan velocity and the faster cooling rate. The base alloy acts as heat absorber (heat sink) and the temperature gradient between the base alloy (substrate) and the incoming laser has significant effects on the microstructures. Else, larger residual stresses may be infused in the microstructures and impart on the surface quality of the com- on the base alloy at 293 K. posite coatings. The spot diameter and the distance between the base alloy and the laser nozzle influences the peak temperature distribution in the molten pool. The peak temperature at the start was 529 K while at the end of the base alloy it increased to 534 K. The crystal structures in the melt pool are changed by the rate of solidification as shown in Figs. 8 and 9. The increase in the scanning speed of the LMD operation produced less dense composites, there was not enough time for the coating materials to interact with the Laser and fuse with the base alloy. The increase in the scanning speed of the LMD operation decreased the deposit width, while the decrease in scanning speed increases the deposit width; resulting from the available time for the coating materials to interact with the laser and fuse with the base alloy [34,35]. From the temperature distribution results, it is evident that the increase in the scanning speed of the LMD operation decreases the rate of dilution of the coating materials within the structural matrix of the base alloy [35]. The propagation of these grains were initiated by the solidification process of the microstructure after the LMD operation. The laser interacts with the powders (coating materials) fed onto the surface via the feed nozzle, the laser melts the powder coatings before they arrive on the surface of the base alloy (Ti-6Al-4 V) creating a melt pool of the deposited materials directly on the surface of the base alloy. The changes in the temperature shows the interaction in the melt pool between the laser and the coating materials as the laser moves from its original position towards the end of the base alloy (substrate). Only the region of interaction between the melt pool of the coating materials and the base alloy undergo a microstructural evolution. Hence, the rate of dilution of the composites was required to understand the level of diffusion of the coating materials in the structural matrix of the base alloy. The alloying metals Al and Cu formed dendrites after the solidification process of the composites was initiated as shown in Figs. 8 and 9. Thermal gradient is the major factor that determines if dendrites would be formed in a microstructure. The alloying metal Cu from microstructural studies is known to possesses strong Beta (b) stabilizing particles that are propagated within the microstructural matrix of composites [36\u201338]. Hence the fusion of the Cu alloying element with the Titanium structural matrix initiates the propagation of the Beta-Titanium (b-Ti) phase structure. The Binary alloy Al-Cu within the structural lattice typically propagate dendritic grain structures at increased laser power and scanning speed [39\u201341]. The combination of the alpha phase particles and the beta phase stabilizers form a network of intermetallic inter-dendritic eutectic phase structure within the fusion regions of the composites [42\u201344]."
        ]
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure59.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure59.1-1.png",
        "caption": "Fig. 59.1 Concept 1 (left) and Concept 2 (right)",
        "texts": [
            " \u2022 It should be possible for the robot be used also, as an attachment to an existing vacuum cleaner; in that case, need for separate vacuum generator could be eliminated, and material could be disposed to dust chamber of the main cleaner. \u2022 Need for batteries could be eliminated using direct powering, power from main vacuum cleaner, etc. \u2022 Reaching the space is attained using a combination of sensors and a path algorithm. Concept 1: The foldable aircraft or three-point support structure The proposed design is of a reconfigurable robot which could be used for cleaning flat surfaces. The robot can be operated in two configurations (Fig. 59.1 left): the expanded configuration and the contracted configuration. The basic components include two cuboidal structures connected to a vacuum. Each cuboidal structure will host its basic circuitry as well as a straight and a rotary brush. The straight brushes will push the dust towards the rotary brushes thereby enabling the rotary brushes to sweep the dust towards the suction. The vacuum will use suction to clean the area and transport it into a dirt bag. i. The Expanded configuration:Useful for cleaning larger areas",
            " Looking at the problems we sought to solve, this robot can access narrow spaces using its contracted configuration. Depending on the boundaries of the area to be cleaned, it can reach the entire area either by expanding or by contracting. Reconfigurability of the robot ensures faster coverage of the area to be cleaned. Addition of a camera to send in live feed to the user or automated image processing could help detect the area to be cleaned. The use of straight and vertical brush as well as suction ensures deep cleaning of the surfaces. Concept 2. The worm-inspired reconfigurable design This design (Fig. 59.1 right) also proposes a reconfigurable robot. It has two configurations: the expanded configuration and the contracted configuration. The basic structure consists of three components: two cuboidal structures hosting internal circuitry and an actuator-suction component connected via tubes. It is integratedwith a straight brush and a vertical or rotary brush for effective cleaning. The straight brushes are added specifically to clean edges. The dust brushed off is pushed towards the suction by the rotary brushed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure12-1.png",
        "caption": "Fig. 12. On-load field distribution of CPM-II with (a) \u03b4=15 deg, (b) \u03b4=30 deg,",
        "texts": [],
        "surrounding_texts": [
            "1) Air-gap flux density Fig. 9 shows the variation of fundamental airgap flux density of CPM-II with slotless stator. The airgap flux density first increases and then decreases as \u03b1 increases. The optimal \u03b1 is about 0.7, which is larger than that of CPM-I. In addition, the variation tendency of flux density with \u03b4 is similar to that with \u03b1, and the optimal \u03b4 is about 60 deg. Fig. 10 shows the fundamental air-gap flux density variation of CPM-II with variable \u03b4. It can be observed that when \u03b4=60 deg, the maximum fundamental airgap flux density can be achieved, meaning that the slotting effect cannot affect the variation tendency of flux density. 2) Magnetic Saturation The open-circuit and on-load field distributions of the CPM-II are shown in Figs. 11 and 12, respectively. It can be observed that the flux density in Region-II is significantly reduced and the saturation of the salient iron pole is weakened as \u03b4 increases. Hence, an excellent diffluence effect can be achieved, and the above theoretical analysis is verified. Figs. 13 and 14 show the on-load magnetic field distribution of CPM-I and CPM-II with \u03b4=90deg under different currents. It can be noticed that saturation of salient core of CPM-I is significant. However, the saturation of salient core of CPM-II with \u03b4=90deg is almost eliminated. (a) \u03b4 Fh Fhy Fhx 0 0 Fhx Fhy r r \u03b4 \u03b4\u03b4 lw Fhx-max Fhy-max PM1 PM2PM2 Stator Air-gap \u03b41 \u03b42 Leakage flux \u03b41 \u03b42 Authorized licensed use limited to: Rutgers University. Downloaded on May 20,2021 at 12:02:13 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (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. (c) \u03b4=45 deg, (d) \u03b4=60 deg, (e) \u03b4=75 deg, (f) \u03b4=90 deg, (g) \u03b4=105 deg, and (h) \u03b4=120 deg."
        ]
    },
    {
        "image_filename": "designv11_63_0000993_s106879982001002x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000993_s106879982001002x-Figure1-1.png",
        "caption": "Fig. 1. Transmission elements of the tail rotor: (1) flange; (2) thrust coupling; (3) fastener; (4) tail shaft tube; (5) support; (6) frame; (7) adapter flange; (8) rubber support; (9) ball bearing; (10) lock ring, (11) nut; (12) rubber seals; (13) body; (14) adapter flange; (15) nosepiece; (16) cap; (17) spline connection.",
        "texts": [
            " 1 2020 8 The tail rotor of the Mi-26 helicopter consists of a shaft line and two gearboxes, namely, the intermediate gearbox and tail rotor gearbox [8]. Shaft line consists of nine tubular shafts (seven\u2014in the horizontal fuselage boom, two\u2014in the fin), connected by seven gear spline couplings, one part of which (nosepiece) rotates in a bearing fixed by means of a mounting flange on the frame. In this case, six couplings are installed on the fuselage frames, and one\u2014 on the rib of the fin. Elements of the horizontal part of the shafting are shown in Fig. 1. According to the results of the analysis performed, the operating time of the gear couplings of the tail shaft from the beginning of operation to rejection (lateral clearance, tooth runout, corrosion) varies within very wide limits\u2014from 111 to 2985 hours, i.e. coupling failure is possible in any period of their operation. Therefore, trouble-free operation of the helicopter is achieved by regular monitoring the state of the couplings and bearings of supports every 100 hours of flight or at least once every 12 months"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure14-1.png",
        "caption": "Fig. 14 Dynamic analysis in configurations 1 and 3",
        "texts": [
            " Assuming that the active part 1 rotates at a constant speed of 6 r/min (motion period is 10\u00a0s), the dynamic simulation is carried out in SolidWorks virtual prototype environment, and the relationship between the driving torque and time of the planar double-folded metamorphic mechanism is obtained. When the planar double-folded metamorphic mechanism is in configurations 1 and 3, the mechanism under force constraint can be regarded as consisting of an active part and an augmented Assur group RRRP (as shown in Fig.\u00a012b and c). The dynamic analysis of the planar double-folded metamorphic mechanism in configurations 1 and 3 is shown in Fig.\u00a014. The dynamic equations in configurations 1 and 3 can be obtained by Eqs.\u00a0(1) and (4) as follows, When the mechanism is in configuration 2, the mechanism under geometric constraint can be regarded as consisting of an active part and an Assur group RRR, as shown in Fig.\u00a015. The dynamic analysis of the planar double-folded metamorphic mechanism in configuration 2 is shown in Fig.\u00a014a and b. The dynamic equations in configuration 2 can be obtained by Eqs.\u00a0(1) and (6) as follows, (11) \u23a7\u23aa\u23a8\u23aa\u23a9 Fax + Fbx = m1a1x Fay + Fby \u2212 G1 = m1a1y 0.5LABFbx sin(\ud835\udf031 \u2212 \u03c0) + 0.5LABFby cos(\ud835\udf031 \u2212 \u03c0) +M1 \u2212 0.5LABFax sin(\ud835\udf031 \u2212 \u03c0) \u2212 0.5LABFay cos(\ud835\udf031 \u2212 \u03c0) = J1?\u0308?1 (12) \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23a9 \u2212Fbx + Fdx = m2a2x + m3a3x \u2212Fby + Fdy \u2212 G2 \u2212 G3 = m2a2y + m3a3y \u2212LBCFbx sin \ud835\udf032 + LBCFby cos \ud835\udf032 + 0.5LBCG2 cos \ud835\udf032 = JBC?\u0308?2 \u2212LCDFdx sin \ud835\udf033 + LCDFdy cos \ud835\udf033 \u2212 G3l3 cos(\ud835\udf033 + \ud835\udefc3) +Md = JDC?\u0308?3 \u2212Fdx = m4a4x \u2212Fdy + Fe \u2212 G4 = m4a4y l4Fdx +Me \u2212Md = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure44.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure44.5-1.png",
        "caption": "Fig. 44.5 Front view of a line diagram of basket depicts pivot point from base",
        "texts": [
            " The main basket is with height 57 mm developed with reference to the smallest glass that we considered in our range, which is a paper teacup with a height of 64 mm. Thus, even the shortest cup is significantly accessible. Importantly, the pivot point should be higher than the center of gravity in order to maintain equilibrium. The tallest glass considered is juice glass with a height of 110 mm. Considering the CG for the tallest glass, which is 65 mm from the bottom as calculated we need to shift the pivot 10 mm higher to maintain equilibrium (Fig. 44.5). Ta bl e 44 .2 C on ce pt s co m pa ri so n F ig .2 .1 C on ce pt -1 F ig .2 .2 C on ce pt -2 (m et al ) F ig .2 .3 C on ce pt -3 F ig .2 .4 C on ce pt -4 \u2022 Si m pl e fo ur -p oi nt lo ck in g de si gn \u2022 E as y to at ta ch an d de ta ch \u2022 A bl e to ho ld gl as se s an d cu ps as w el l \u2022 L ig ht w ei gh t( 25 .3 g) \u2022 M at er ia lu se d is A B S PC \u2022 It ca n be ar a w ei gh to f 22 0 g \u2022 M et al us ed fo r st re ng th an d ri gi di ty \u2022 U se d m at er ia li s an ne al ed ca rb on st ee l( SS ) PC (f oo d gr ad e m at er ia l) \u2022 It is al m os ts ev en tim es he av ie r th an A B S PC 18 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001018_isef45929.2019.9097061-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001018_isef45929.2019.9097061-Figure2-1.png",
        "caption": "Fig. 2. Structure of proposed pole.",
        "texts": [
            " Therefore, in this paper, the torque in the full model is obtained by summing up torques of the simple magnetic pole model. TABLE I. SPECIFICATIONS OF SPHERICAL ACTUATOR Outer diameter of rotor [mm] 97.0 Outer diameter of stator [mm] 82.0 Residual magnetic flux density [T] 0.68 Number of coil turns 180 Movable range Around X and Y axes [deg] 40 Around Z axis [deg] 360 (continuously) Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on July 27,2020 at 02:36:03 UTC from IEEE Xplore. Restrictions apply. III. STRUCTURE AND TORQUE CHARACTERISTICS The proposed pole structure is shown in Fig. 2. The conventional pole structure is a simple structure which a coil was wound around a cylindrical iron core. On the other hand, as shown in Fig. 2, an auxiliary pole structure is added around the coil in the proposed pole structure. In the proposed pole structure, a magnetic flux make a loop between the main and auxiliary poles via the back yoke. The torque characteristics of the conventional and proposed models are compared using FEM, and the effectiveness of the proposed model is investigated. In this paper, as an example, the torque characteristics when the simple poles are rotated in the latitudinal direction are evaluated. The poles face the center of the permanent magnet at a longitude of 0 deg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002961_j.prostr.2021.03.004-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002961_j.prostr.2021.03.004-Figure2-1.png",
        "caption": "Fig. 2. Shaker table test \u2013 SLM bracket assembly parts 94 mm",
        "texts": [
            " The minimum \u201820g\u2019 capability was required for the SLM bracket to cope with the vibrations from other prime sources during engine normal operation and wind-milling vibrations post an in-flight shutdown of the engine, may be due to an ultimate failure event. The optimized AM bracket mass was 91.7 grams which was significantly less (> 50%) than the mass of an equivalent sheet metal Ti-6Al-4V bracket (200 grams). The configuration and mounting of the bracket assembly for the shaker test are shown in Fig. 2. A dummy aluminium block was machined to represent the sensor unit and was integrated with the bracket using two 0.25\u201d  bolts. The total mass of the bracket assembly was 306.7 grams (Bracket mass = 91.7 grams, dummy block mass = 179 grams and fasteners mass (2 off) = 36 grams). An additional aluminium fixture was machined as an interface between the SLM bracket and the shaker table. The bracket was attached to the interface fixture via three Inconel 718 0.25\u201d  bolts tightened to a torque of 15.2 N"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001702_5.0016147-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001702_5.0016147-Figure1-1.png",
        "caption": "FIGURE 1. The direction of the load given to the specimen",
        "texts": [],
        "surrounding_texts": [
            "a. The test uses a machine Universal Testing Machine (UTM), and the tests carried out are: 1. The tensile strength of the fibers perpendicular to the y and z axes. 2. Shear strength with y and z axis direction. 3. Compressive strength with x, y, and z axis direction. b. Thickness of bamboo slats."
        ]
    },
    {
        "image_filename": "designv11_63_0001023_iccre49379.2020.9096444-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001023_iccre49379.2020.9096444-Figure3-1.png",
        "caption": "Figure 3 Model of the prototype",
        "texts": [
            " Apart from these three sections another three motors have to be used in order regulate the angle, direction and height of the shuttlecock ejecting point. Then it will have the capability to produce different shots. The arrangement of the required 3-DOF for this, was discussed in the section 02. From now on this arrangement will refer as the shuttle positioning section or direction controlling section. Adhering to these technical and other factors a model was developed using SolidWorks software platform. This model of the prototype is shown in Fig. 3. 75 Authorized licensed use limited to: Murdoch University. Downloaded on June 13,2020 at 23:52:11 UTC from IEEE Xplore. Restrictions apply. One of the project goals was to minimize the possible shuttlecock damage during the operation of the machine. The diameter of the shuttlecock cork base varies from 25mm to 26mm [18]. However, for the ejecting purpose, elastic property of the cork base has to be used. Considering the high initial linear velocity of the rollers relative to the shuttlecock, instant gripping action with the cork base will cause damages to the shuttlecock"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001877_eircon51178.2020.9254082-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001877_eircon51178.2020.9254082-Figure4-1.png",
        "caption": "Fig. 4. Exoskeleton trajectory model obtained (Isometric view 3D).",
        "texts": [
            " Video processing data obtained from an image recognition algorithm allows to get 3 angles associated with a certain time lapse (Fig. 2b) as it is explained in [2]. This data and the environment conditions, like gravity, were used as input parameters for the computational model. The data of angles associated with each DOF is processed by Computer Aided Engineering (CAE) and Finite Element Analysis (FEA) software, as explained in [6,7]. We obtained output parameters such as accelerations, velocities, moments and other useful results, and the modeled trajectory (Fig. 4, Fig. 6) will be used for future analysis related to biomechanical study, design and validation methods, as described in [2,6]. After considering all conditions and input parameters we could transform the base arm trajectory, which has a duration of 3.6 seconds (Fig. 1). This was used for simulation by using software resources as shown in Fig. 4. Once the angles were obtained, this data was passed to CAE software [10] and we got the path traveled by the user arm as shown in Fig. 5; these results were based on practical methodology developed in [2]. It was possible to follow the plane of the wrist and hand trajectory, as Fig. 7 and Fig. 8 shows. Authorized licensed use limited to: American University of Beirut. Downloaded on May 20,2021 at 20:20:52 UTC from IEEE Xplore. Restrictions apply. From the movement model seen in Fig. 5, an arm trajectory was modeled as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002533_062088-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002533_062088-Figure1-1.png",
        "caption": "Figure 1. Phases of the process of passing the heap through the hole of the grating surface: 1 \u2013 a scraper; 2 \u2013 a grating bottom; 3 \u2013 a longitudinal hole; 4 \u2013 elementary volume of a heap; 5 \u2013 a hatchelled heap; 6, 8, 10 \u2013 the second and subsequent layers of the heap; 7 \u2013 separation surface; 9 \u2013 the initial position of the hatchelled heap.",
        "texts": [
            " The main purpose of the theoretical part of the study was to simulate the process of separating free grains from a hatchelled heap using the grating bottom of the inclined chamber of the harvester and a moving mesh surface mounted in front of the threshing drum. In the version with an inclined chamber, the research hypothesis considered processes occurring in the elementary volume of a heap formed by each scraper of a floating conveyor. The separation is as follows. When the heap reaches the edge, the grain openings roll down in layers along the separation line between the inclined layers, oriented at an angle of internal friction of the material to the horizon \u03b1 and pass through the oblong holes of the grating bottom of the inclined chamber (figure 1). At the same time, part of the grain of the moving layer slides along the interface without turning, and part is rolled, accelerating as it moves. If the length of the rolling layer exceeds the path traveled by the grain that was in the initial state in the upper part of the elementary volume during the movement of the scraper above the longitudinal hole of the grating bottom, then its passage is impossible. In this regard, it can occur at one of the following holes located on the path of the scraper and the elementary volume of the heap moved by it"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000834_crc.2019.00024-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000834_crc.2019.00024-Figure3-1.png",
        "caption": "Fig. 3. Model of the wheelchair and robot",
        "texts": [
            " The heights of steps located at the entrances of typical buildings and other structures were also measured, and the target step height was set at 120 mm, which accounts for more than 80% of the observed heights. This paper is organized as follows. Section II describes the cooperative system, and Section III describes the process of climbing a step. Section IV describes a verification experiment and its results, and Section V provides conclusions. The robot used in this research was \u201cTateyama,\u201d a wheeled robot developed in our laboratory (Fig. 3). Table I lists the specifications of the robot. The vehicles are deployed in a forward-and-aft configuration for ascending steps. When the wheelchair and the robot encounter a step, the robot hand grasps the rotary shaft of the push handle of the wheelchair [18]. The robot has three pairs of left and right wheels (Fig. 4). The front pair are casters, and the middle and rear pairs are driving wheels. The arms to which the front and rear wheels are mounted can be deployed and retracted (Fig. 5), and these mechanisms are used in the step climbing process. The robot has manipulators attached to the left and right joints of its upper half; each arm has 5 degrees of freedom (DOFs) and the hand has 1 DOF, for a total of 6 DOFs (Fig. 6). As configured for this study, the length of the upper arm link [from Joint 2 (shoulder) to Joint 4 (elbow)] was l2 and the length of the forearm link (from Joint 4 to Joint 6) was l4C. The manipulator joint angles were \u221290 deg \u2264 \u03c62 \u2264 +90 deg and 0 \u2264 \u03c63\u2264 +100 deg (Fig. 3). The right and left hands each had two fingers (Joint 6) to hold items (Fig. 6). Touch sensors were set on both forearm links, each of which were covered by bumpers (Fig. 7). These touch sensors were used when the robot front wheels climb a step (see Section III). The robot had a stopper mounted on its front of its body (Fig. 8 (a)). The stopper limited the passive rotational travel of the manipulators during wheelchair pushing (Fig. 8 (b)) and enabled the robot to imitate the operation of a human pushing an object (Fig",
            " Both vehicles stop and the robot rear wheels are retracted (folded upward) (Fig. 5). We experimented with the proposed system under an environment with a step 120 mm high and a friction coefficient \u03bc = 0.72 (the floor and step material were wood; Fig. 16). The wheelchair user was an able-bodied adult male (age 20, weight 50 kg), and the experiment was done on one floor of the National Institute of Technology, Toyama College. The velocities of both vehicles were constant (0.76 km/h). In stage 1, the wheelbase between the robot front and middle wheels, WB f (Fig. 3), was set as WB f =400 mm. The system detected the step, the robot stopped, and the wheelchair moved forward. After the front wheels of the wheelchair were 79 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 06:18:12 UTC from IEEE Xplore. Restrictions apply. lifted, both vehicles moved forward, the system was able to put the front wheels of the wheelchair on the step. In stage 2, both the vehicles continued to move forward. The back wheels of the wheelchair come into contact with the step, and the robot continued to push the wheelchair. The robot prevented the wheelchair from tipping over backward, and it supported the ascent of the wheelchair rear wheels. After rear wheels of wheelchair reached the step, the level sensor on the chair detected the end of wheelchair\u2019s ascent. In stage 3, the wheelbase of the robot was changed to WB f = 220 mm (Fig. 3) to achieve the configuration needed for robot climbing. The mechanisms of the front and rear wheels were retracted (Fig. 5), The arms are spread apart, the touch sensors on the forearm links detected contact with the stopper of the wheelchair (Fig. 7), and the dual manipulators were set in the wheelchair stopper (Figs. 11 (a) and (b)). The robot detected the position for lifting the robot front wheels. The wheelchair maintained its position as the robot moved forward. The accelerometer on the robot detected the robot angle for step climbing, and the robot front wheels touched the step"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure6-1.png",
        "caption": "Fig. 6: 3D-modal analysis model",
        "texts": [
            " The 1187 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 16,2021 at 18:43:55 UTC from IEEE Xplore. Restrictions apply. higher the scalar product ~\u03a8\u2032T0,i ~\u0302 f(\u03c9r), the better is the match of force shape and mode shape. The force is exciting the structure with the frequency \u03c9r. Besides analytical calculations [7], [18], [19], a numerical FEA is widely used to predict the eigenfrequencies. The structural dynamics of the machine were studied by a 3D-modal analysis in ANSYS Mechanical (fig. 6) and by an experimental modal analysis on the specimen. The FE-model consists of the stator, coils in the stator slots and an aluminum housing. In addition, the end shields and the flange were considered by an elastic support boundary condition indicated by the blue color in the screw threads of the housing. Moreover, this boundary condition allows torsional mode shapes to occur in the simulation. The experimental setup is shown in Fig. 7. Triaxial accelerometers were placed on the outer housing surface of the machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002004_ddcls49620.2020.9275203-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002004_ddcls49620.2020.9275203-Figure2-1.png",
        "caption": "Fig. 2: Right wing frame, right stroke plane and body frame",
        "texts": [
            " For stroke plane frames, they are located at each hinge point for right or left wing, their orientation parallel to the B frame at the initial time and its origin at the wing joints. Furthermore, the stroke plane frames are obtained by rotating around the second coordinate axis of the B frame, hence they have only one degree of freedom with respect to it, the stroke plane angle \u03b2. For left and right wings, the orientation of z\u2212 axis of stroke plane frames are opposite. So the rotation matrix between right and left stroke plane frame andB frame and are as (2) and (3), respectively. The right wing frame, right stroke plane and body frame are shown in Figure 2. RBI = CB2C21C1I =  1 0 0 0 cos(\u03c6) sin(\u03c6) 0 \u2212 sin(\u03c6) cos(\u03c6)  cos(\u03b8) 0 \u2212 sin(\u03b8) 0 1 0 sin(\u03b8) 0 cos(\u03b8)   cos(\u03c8) sin(\u03c8) 0 \u2212 sin(\u03c8) cos(\u03c8) 0 0 0 1  (1) RSRB =  cos (\u03b2R) 0 \u2212 sin (\u03b2R) 0 1 0 sin (\u03b2R) 0 cos (\u03b2R)  (2) RSLB =  cos (\u03b2L) 0 \u2212 sin (\u03b2L) 0 1 0 sin (\u03b2L) 0 cos (\u03b2L)  1 0 0 0 \u22121 0 0 0 \u22121  (3) The third reference frames are the wing frames, which also include into right and left frames. The origin of each wing frame is the joint between the wing root and the body. Other than that, the x\u2212 axis and y\u2212 axis of each wing frame are initially defined in the stroke plane and the z\u2212axis is downward or upward and orthogonal to this plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001242_tec.2020.3007715-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001242_tec.2020.3007715-Figure11-1.png",
        "caption": "Fig. 11: Winding excitation connection for (a) stator and (b) rotor.",
        "texts": [],
        "surrounding_texts": [
            "For the measured load performance, the prototyped test motor is synchronized with the 400 V grid supply and loaded by varying the dynamometer power in the test bench of Fig. 12. For these tests, the field current was set constant I f = 23 A and the winding temperature monitored at 75 \u25e6C. The measured load performance results as a function of input power are shown in Fig. 15. Authorized licensed use limited to: University of Exeter. Downloaded on July 16,2020 at 00:15:25 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. To predict the load performance, the iterative process of Fig. 6 is used in which the grid load angle is calculated using the mapping technique given in Section V. The performance curves of power factor, efficiency, power input, power output and current are predicted by (12), (14), (15), (20) and (27) respectively. The predicted load performance curves as a function of input power are shown in Fig. 15. This shows good agreement between predicted and measured load performance. The slight difference in the power factor and current in Fig. 15(b) at low input powers can be explained by the FEM used BH curve. This aspect is further considered in Sub-section VIII-A."
        ]
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure1-1.png",
        "caption": "FIGURE 1. Model of electric bus\u2019s frame",
        "texts": [
            " The purpose of their project is to redesign the parameters like gauges, materials, geometry/shapes, topology variables, weld pitch, joints for improving the structure so that the total weight of the bus is reduced. Satrio Wicaksono et al. [11] did the bus roll over test in accordance with UN ECE R66 Standard. Investigation of the bus in empty and fully loaded condition showed that the superstructure of the bus is not strong enough, as the residual space safety criterion was violated. The model of bus frame has been draw using Solidworks Software, the model was depicted in Figure 1. The detail dimension of electric bus\u2019s part is tabulated in Table 1. Three type of material were used in the simulation, namely: JIS 3445 STKM 13A, Aluminum 6005A T6 and Aluminum 6061 T6. The properties of material were shown in Table 2. 030153-2 Boundary Condition (BC) must be set on the model before simulation. BC must represent the real condition of object. In this paper BC was lied on the bottom part of frame since this part was connected to the chassis of electric bus. The type of BC is fixed which the rotation and displacement are not allowed in all directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002741_033080-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002741_033080-Figure1-1.png",
        "caption": "Figure 1. The main components of an electric bicycle.",
        "texts": [
            " In this paper, the process of increasing the speed setting performance 350-Watt BLDC motors used in electric bicycles, through mathematical modelling and optimization. The optimization method carried out in this study is by PID controller and hybrid fuzzy logic. The identified results used to determine the performance of the BLDC motor speed control are indicators of transient response. The sign of transient response is to be measured by the value of peak amplitude, rise time, and settling time. In Figure 1 describe some of the main components of an electric bicycle, such as a battery, controller, throttle, and BLDC motorbike. The BLDC motor used in this study has a maximum power of 350 watts. Data input data includes pulse width modulation (PWM), voltage, and current. While the output data is in the form of motor rotational speed, input and output data used to identify the BLDC motor system, to obtain a mathematical model in the way of a transfer function. After getting a mathematical model in the form of a transfer function, optimization is carried out by using a PID controller, fuzzy logic and hybrid fuzzy logic PID controller Evaluation results to determine the performance of setting the BLDC motor speed then simulated using the MATLAB program for transient response values such as peak amplitude, rise time, and settling time Mathematical modelling of BLDC motors is done through an experimental process to obtain input and output data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002216_s12541-020-00457-y-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002216_s12541-020-00457-y-Figure13-1.png",
        "caption": "Fig. 13 Various shapes of compliant link 3 and the result of stress simulation",
        "texts": [
            " If the material is more compliant (less stiff), it can only endure small external loads. The task was to design link 3 as compliant as possible; however, the link should not exceed yield stress, which leads to fracture. A compromise between stiffness and strength was made when designing compliant link 3. To design a single-body link 3 with such values, a series of steps were taken. First, the material was chosen to be SUP9 spring steel just as previous study [27]. Next, various spring-shaped design alternatives were considered. Figure\u00a013 shows the design alternatives with different shapes. For each design alternative, several derivations with small dimension changes were made. Consequently, a total of 69 candidates of link 3 were generated. For all link 3 candidates, both the compliance and the maximum 1 3 stress were calculated. The simulation was performed by the CAD software SolidWorks. The colors in Fig.\u00a013 show how stress is distributed in each design alternative. The graph of k and the maximum stress of all 69 candidates is presented in Fig.\u00a014. As shown in the figure, it is difficult for candidates to have both small k value and small maximum stress. With smaller values of k, there is a tendency of having a larger maximum stress. This tendency forms the frontier boundary in the graph which is called the \u201cpareto front.\u201d The candidate displayed by the red dot in Fig.\u00a014 was chosen to be the final compliant link 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001408_1077546320945472-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001408_1077546320945472-Figure2-1.png",
        "caption": "Figure 2. One-degree-of-freedom flexible manipulator.",
        "texts": [
            " (2015) developed a BTS with a flexible-link slave manipulator, where the tip position tracking problem was addressed by the output redefinition method. Rasouli et al. (2020) considered a BTS composed of a rigid master robot and a flexible-link slave robot; they proposed a simple PD control structure to handle the control aspect of the system in the presence of actuator faults and disturbances. In many recent studies, BTSs are lumped parameter systems described by ordinary differential equations (ODEs). In this study, a BTS composed of master-slave manipulators with link flexibility is considered (see Figure 2). Given that the deformation and vibration of the flexible manipulators are related to position and time, the BTS is essentially a distributed parameter system (DPS). Common ODE models ignore the higher-order modes of the DPS, whichmay initiate a control spillover problem. However, a partial differential equation (PDE)-based dynamic model can retain all modes to restrain spillover and accurately describe the DPS. Xing et al. (2019) established a PDE model for a flexible string system. A flexible aerial refueling house was modeled as a DPS expressed by PDEs (Liu et al",
            " \u00f0\u2217\u00de0 \u00bc \u2202\u00f0\u2217\u00de=\u2202x; \u00f0\u2217\u00de00 \u00bc \u22022\u00f0\u2217\u00de=\u2202x2, \u00f0\u2217\u00de000 \u00bc \u22023\u00f0\u2217\u00de=\u2202x3, and \u00f0\u2217\u00de\u00f04\u00de \u00bc \u22024\u00f0\u2217\u00de=\u2202x4 represent the partial derivatives of \u00f0\u2217\u00de with respect to space. We consider the BTS shown in Figure 1. The human operator and the flexible master manipulator are in the control room, whereas the flexible slave manipulator is in the remote environment. The trace information and control inputs are transmitted through the control and communication medium. We assume that there is no communication delay in the BTS. As shown in Figure 2, themaster and slave robots use the same flexible manipulator with one DOF. In Figure 2, XOY is the global coordinate system and xOy represents the bodyfixed coordinate system. System parameters are defined as follows: each flexible master manipulator and slave manipulator has length L, hub inertia Ih, bending stiffness EI , mass per unit length \u03c1; and payload mass m. Fm(t) and Fs(t) denote the force inputs generated by actuators. Mm(t) and Ms(t) represent the torque inputs generated by motors. \u03b8m\u00f0t\u00de and \u03b8s\u00f0t\u00de are joint angles, ym\u00f0xm; t\u00de and ys\u00f0xs; t\u00de denote the elastic deflections, wm\u00f0xm; t\u00de \u00bc xm\u03b8m \u00fe ym\u00f0xm; t\u00de and ws\u00f0xs; t\u00de \u00bc xs\u03b8s \u00fe ys\u00f0xs; t\u00de represent the displacements of the flexible master\u2013slave manipulators, and it easily follows that w\u00f0\u03c5\u00de m \u00f0xm; t\u00de\u00bc y\u00f0\u03c5\u00dem \u00f0xm; t\u00de and w\u00f0\u03c5\u00de s \u00f0xs; t\u00de\u00bc y\u00f0\u03c5\u00des \u00f0xs; t\u00de hold when \u03c5 \u2265 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001573_j.ijsolstr.2020.09.004-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001573_j.ijsolstr.2020.09.004-Figure7-1.png",
        "caption": "Fig. 7. Modeling of compliant boundary condition with three longitudinal springs in three direction and one rotation spring normal to the plane.",
        "texts": [
            " 10 and 11, where the confined corridor and alternating shapes are obtained by the summation of low-energy in-plane compliant modes. Ultimately, Table 1 shows the value of parameter d3 and two other geometrical parameters, d1 2 to ensure a lattice with the minimum required in-plane actuation force. The boundary conditions are designed to have minimum effect on the in-plane movement of the kagome lattice under large displacement values. This is accomplished by using leaf springs at boundary points as compliant boundary conditions, see Fig. 7. The leaf springs are designed from the same grade of steel as the structure. Compliant boundary conditions are modelled as three longitudinal springs Ki i \u00bc 1;2;3\u00f0 \u00de and one torsional spring K6. Denoting geometrical parameters of the leaf spring with S1 3, as shown in Fig. 7, the equivalent stiffness in span (K1) and chord (K2) directions is calculated based on the deflection of a cantilever beam with zero slope at the tip point, where its movement is confined within the plan of the kagome lattice. The equivalent stiffness in normal direction (K3) is determined as the stiffness of bar element with the cross-sectional area of S1S2. The torsional stiffness K6 is estimated based on the torsion constant J given in Young and Budynas (2002) for bars with rectangular cross section",
            " Importantly, the tightening of internal members causes a significantly large increase in actuation force, as shown in Fig. 8, in comparison with the linear solution that ignores the additional stiffness induced by the stress state. The tightening of internal members also causes a lateral shrinkage deformation toward the center of the kagome lattice structure in Fig. 13(b). To mitigate the tightening problem, some internal members can be removed. This implies the necessity of optimizing the kagome lattice design ing the relations given in Fig. 7. to find an optimum continuous lattice with the aim of reducing actuation force under large deformation. As can be observed in Eqs. (15b) and (15c), the nonlinear stiffness matrix and the internal force vector are determined based on the instantaneous stress field in each increment. By removing each member of the lattice, the stress distribution changes in the new structure which in turn will change the incremental deformation path of the lattice under large amplitude actuation. As per the large numbers of configurations that can be obtained by removing different members of the lattice as well as the nonlinear deformation path which each candidate configuration will undergo, the optimum design is obtained using Bayesian optimization (Frazier, 2018), which is a class of machine-learning-based optimization methods"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002806_tia.2021.3079169-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002806_tia.2021.3079169-Figure2-1.png",
        "caption": "Fig. 2. Surface permanent magnet motor structure in xy cross-sectional view.",
        "texts": [
            " PROPOSED MACHINE STRUCTURE Figure 1 shows a cut view of the fabricated prototype machine of the proposed single-drive bearingless motor by the authors for cooling fan applications [18]. This machine has a single-drive bearingless motor and two sets of repulsive passive magnetic bearings. At the top of the rotor shaft, a fan blade is installed. At the bottom, a displacement sensor is installed to detect the rotor axial displacement. One set of three-phase winding is wound in the center stator core as the concentrated winding. Figure 2 shows the xy cross-sectional view of the singledrive bearingless motor. The cross section is identical with a surface permanent magnet machine so that the rotational torque is generated by q-axis current. The numbers of rotor poles and stator slots are eight and twelve, respectively. The active axial force is generated by d-axis current in the singledrive bearingless motor. The detailed principle of the active axial force generation is shown in [18]. Figures 3 (a) and (b) show principles of passive stabilization in the radial and tilting motions by the repulsive passive magnetic bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002881_s42417-021-00329-3-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002881_s42417-021-00329-3-Figure8-1.png",
        "caption": "Fig. 8 Finite element model of gear transmission system",
        "texts": [
            " In the finite element analysis of structure, each node usually has three degrees of freedom, while for the thermal analysis element, it usually has only one temperature degree of freedom. Because the finite element thermal analysis in this paper is a three-dimensional model, the commonly used three-dimensional thermal analysis elements are introduced as follows: To simulate and analyze the dynamic temperature field, a three-dimensional finite element model of gear transmission (11)[M]{x\u0308(t)} + [C]{x\u0307(t)} + [K(t)]{x(t)} = 0 (12)[M]{x\u0308(t)} + [K(t)]{x(t)} = 0 (13) ( k \u2212 wim ) i = 0 (14)k \u2212 wim = 0 system is established as shown in Fig.\u00a08. Because the analysis of temperature field is needed, solid70 hexahedral 8-node element is used in mesh generation based on the introduction in the previous section. However, the hexahedral element is difficult to be divided in the transition position of the tooth root, so solid87 tetrahedral 10 node element is used to fill. After analysis, it is found that the value of element quality is: 0.82. Therefore, the quality of the model is better. For different components in the system, due to different materials, the corresponding properties are also inconsistent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000947_kem.841.381-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000947_kem.841.381-Figure2-1.png",
        "caption": "Fig. 2 Stress distributions of the coil spring.",
        "texts": [
            " From the manganese and carbon compositions, it can be concluded that the material was included into the low carbon steel. According to ASTM A322-91 [19], the material used to fabricate the coil spring was the SAE 5160 carbon steel. The coil spring material is ideal when a high ultimate tensile strength and yield strength with a low material modulus of elasticity is able to store the maximum energy received [20]. The coil spring stress concentration is represented in red, orange, yellow, green and blue in Fig. 2. The maximum stress was 0.99751 GPa, which occurred inside the coil spring and smaller compared to the outside region. The maximum stress was below the yield strength, i.e. 1.487 GPa [21], which indicated that the load did not cause a failure at the component. The critical point also occurred at the border between active and inactive coils in accordance with the other findings [3, 22-24]. Fig. 3 shows the measured strain signals with a non-zero-mean value, which was -688 \u00b5\u025b, -678 \u00b5\u025b and -714 \u00b5\u025b for the flat, uphill and downhill roads, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002066_tie.2020.3045591-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002066_tie.2020.3045591-Figure5-1.png",
        "caption": "Fig. 5. Virtual voltage vector deviation \u2206V c, auxiliary x-axis and auxiliary angle \u03b3",
        "texts": [
            " From (14), the d and q components of the virtual current response \u2206ic can be shortly described as \u2206icd = Ts Ld \u2206V cd \u2206icq = Ts Lq \u2206V cq (15) where \u2206ic = \u03b4ic2\u2212 \u03b4ic1, \u2206V c = \u03b4V c2\u2212 \u03b4V c 1, the amplitude of \u2206V c is \u2206V c. The main difference with (11) is that in (15) virtual variables are used and the estimation of steady-state voltage vector V is not necessary. 3) Pixel points of images in auxiliary xy-reference frame: The auxiliary xy-reference frame is introduced, and the deformed and primitive images are generated in this auxiliary reference frame. As shown in Fig. 5, the introduced x-axis aligns with the direction of the virtual voltage vector deviation \u2206V c. The auxiliary angle \u03b3 is the direction of the introduced x-axis with respect to the d-axis. The information of the virtual voltage vector deviation \u2206V c in the \u03b1\u03b2-reference frame, including the angle \u03b3\u2032 and the amplitude \u2206V c, can be derived from applied inverter voltage vectors. For instance, as shown in Fig. 3, in the case that the inverter voltage vectors sequence is V 5\u2192V 3, the angle \u03b3\u2032 = \u03c0 2 , and the amplitude \u2206V c = \u221a 3 2 2 3Vdc = \u221a 3 3 Vdc. As shown in Fig. 5, the introduced auxiliary angle \u03b3 can be used to project the variables along d- and q-axes to x- and y-axes. By rotational transformation of dq components to xy components, equations (15) can be written as \u2206icx = Ts Ld \u2206V c cos2 \u03b3 + Ts Lq \u2206V c sin2 \u03b3 \u2206icy = \u2212 Ts Ld \u2206V c sin \u03b3 cos \u03b3 + Ts Lq \u2206V c sin \u03b3 cos \u03b3 (16) Equations (16) can be rewritten as \u2206icx Ts\u2206V c = cos2 \u03b3 Ld + sin2 \u03b3 Lq \u2206icy Ts\u2206V c = \u2212 sin \u03b3 cos \u03b3 Ld + sin \u03b3 cos \u03b3 Lq (17) where the left-hand side indicates deformed pixel points ( \u2206icx Ts\u2206V c , \u2206icy Ts\u2206V c ), and these points are based on virtual current responses in the auxiliary xy-reference frame, the DC-link voltage and the inverter voltage vectors sequence",
            " The magnetic saliency aligns with the geometric pole of the rotor d-axis in the case that the cross-saturation effect can be neglected. When the machine operates under no load, the magnetic saliency aligns with the geometric pole of the rotor, as shown in Fig. 8(a). When the magnetic state changes, such as the machine operates under a rated load, the cross-saturation can not be ignored, and the magnetic saliency rotates to ds-axis accordingly, as shown in Fig. 8(b). The self-sensing estimation is consequently subject to an error \u2206\u03b8. Based on Fig. 5, the estimation error of auxiliary angle \u03b3\u0303 has the same modulus of the estimation error of rotor position \u03b8\u0303. Considering the crosssaturation effect only, the location of the deformed point is rotated a certain angle 2\u2206\u03b3 from the primitive point along the circle by the cross-saturation effect, as shown in Fig. 9. By using the rotation operation in transformations T\u2217j , the deformed point G can be shifted to transformed point K. Operations of rotation, scale, translation and shear in transformations T\u2217j are used to reduce the location deviations of deformed points caused by several non-ideal factors that distort current responses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000089_ev.2019.8892977-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000089_ev.2019.8892977-Figure2-1.png",
        "caption": "Fig. 2. Topologies of PMVM with magnets in outer stator",
        "texts": [
            " In this section it is presented the comparison design and the influence of the PMs\u2019 position in the outer stator in order to achieve high torque with a consistence of torque ripples as low as possible. All the structures where inspired from [7] - [9]. Considering that the air gap length is doubled, it doesn\u2019t influence the efficiency badly even on the contrary, it keeps it over then 90%. In comparison with the dual stator spoke Vernier machines [9] those structures have passed the threshold of 95% efficiency and output power closer to 2 kW. The proposed structures are represented in Figure 2 with considerable construction differences. PMVM-SMMRS in Figure 2 a) has a similarity for the space between thickness of both stator secondary poles and rotor poles with PMs on interior surface. PMVM-BM in Figure 2 b) the magnets are buried in outer stator and rotor in order to see the influence of the lack of air reluctance. PMVM-SMM in Figure 2 c) the magnets are mounted on surface to improve the uniformity of the iron core. The magnets volume and dimensions are the same for all the proposed structures. For all the structures, red is phase A, yellow is phase B, and black is phase C. The stator core is distributed in two regions, the inner region contains 9 main poles with 3 secondary poles each. a) PMVM with surface magnets mounted in rotor and stator poles (PMVM-SMMRS) The outer region is based on 27 negative and 27 positive magnets oriented radial which gives an advantage in magnetizing the magnetic core"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001068_10402004.2020.1767251-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001068_10402004.2020.1767251-Figure4-1.png",
        "caption": "Fig. 4 Geometric parameters of an involute spur gear",
        "texts": [
            " The objective of the EHL model developed in this paper was to obtain the probability distri- bution of the minimum film thickness at different positions along the line of action under a random-meshing force. The EHL model consists of the Reynolds equation, film thickness equation, viscositypressure equation, densitypressure equation, and force balance equation. In this model, thermal effects are neglected. The equivalent radius of curvature is different along the line of action. According to EHL theory, the contact of two gear teeth is assumed to be two cylinders (see Fig. 4). The radii of the cylinders can be defined as follows[28]: 1 b 1 tan ( )R R s  (11) 2 b2 tan ( )R R s  (12) b 1 b 2 b 1 b 2 R R R R R   (13) where s is the distance of the mesh position to the pitch point and represents the pressure angle. The tooth entrainment velocity can be defined as follows [29]: 1 2 ( ) 2U u u  (14) where 1 1 1 2 2 2 , u R u R   , 1  and 2  are rotation speed of the pinion and gear, respectively. Acc ep te d M an us cr ipt It is assumed that the shaft has no misalignment and the effect of contact edges is negligible[13]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000413_s10846-019-01133-8-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000413_s10846-019-01133-8-Figure2-1.png",
        "caption": "Fig. 2 Load cell location graph for measuring torque",
        "texts": [
            " The biomechanics of human knee joint during normal walking were studied for the design of exoskeleton. The knee exoskeleton system includes a Maxon EC40 motor, an exoskeleton leg and a cable drive transmission where the transmission ratio is 15:1. A torque sensor was installed between the pulley and the exoskeleton arm for torque measurement. The angular of the exoskeleton leg is measured with a Nano-Ahrs measurement unit. A force sensor was installed on the leg brace for the measurement of the force exerted by the wearer leg on the exoskeleton leg, as is shown in Fig. 2. The net torque can be calculated as \u03c4i = Fi \u2022 R, where is the output force of the shear force sensor and is the distance between the sensor and the joint center of the knee exoskeleton. The control algorithm of knee exoskeleton is shown in Fig. 3. It consists of admittance model, inertia compensator, muscle torque generation based on adaptive frequency oscillator. The primary input is the torque of man-machine interaction \u03c4i, and a muscle generated torque \u03c4mg is added to the input as assistive torque compensation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001216_j.anucene.2020.107677-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001216_j.anucene.2020.107677-Figure1-1.png",
        "caption": "Fig. 1. Schematic of centre-return reflux ALIP.",
        "texts": [
            " Electromagnetic pumps (Baker and Tessier, 1987; Blake, 1956; Voldek, 1970) are used mainly in the auxiliary circuits of fast reactors and in experimental facilities (Andreev et al., 1978; Anisimov et al., 2012; Kirillov, 1982). Development of a few large electromagnetic (EM) pumps (specially of Annular Linear Induction Pump (ALIP)) for main heat transport system is also reported in literature (Aizawa et al., 2011; Yang and Kraus, 1977; Kliman, 1979; Ota and Katsuki, 2004). A typical schematic of centre-return ALIP is depicted in Fig. 1 where the 3-phase distributed winding creates a traveling magnetic field. Circulating eddy currents are induced in liquid metal as a result of this traveling magnetic field. Interaction of current and magnetic field leads to creation of pumping force. Flow of liquid metal takes place in an annular channel formed by two austenetic stainless steel pipes. The outer pipe is towards the winding. The inner pipe contains magnetic material to provide a low reluctance path for magnetic flux. The annular gap and total non-magnetic air\u2013gap are two of the important parameters that govern the performance of ALIP (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000162_012003-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000162_012003-Figure4-1.png",
        "caption": "Figure 4. Stress (a) and total displacement (b) distributions of the first analysis of the current model.",
        "texts": [
            " For this, the printed control arm was tested in the buggy car by assembling it on the suspension system. The results of reaction forces in the lower control arm are shown in Figure 3. As expected, the forces in the z-component were the most significant, and the x-components were almost nil, although, all of them were low. The resultant forces used for the first and the second analyses were 68.9 N and 82.6 N, respectively. From the two analyses, the first was the most critical since it had higher stress and displacement values, see Figure 4(a) and Figure 4(b). The equivalent von Mises stress distribution resulted from the first analysis had a maximum value of approximately 22 MPa, and the maximum total displacement was approximately 5E-5 m. As previously mentioned, in order to carry out the process of shape optimization, an oversized control arm was designed to later remove the zones that presented low stresses. Figure 5 shows the equivalent von Mises stress of this oversized part, the maximum value was near 3 MPa. To achieve the optimal shape, it was necessary to perform 29 iterations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure44.4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure44.4-1.png",
        "caption": "Fig. 44.4 Glame",
        "texts": [
            " It is important to accommodate all types of glass and cups, which it does with an adjustable base (Fig. 44.3; Table 44.2). This journeywas started in order to cover all the facets of the design considerations, and concept 4was foundmost compatiblewith all aspects of the design consideration; hence, the journey comes to an end with the selection of the concept 4. Design enfolds sturdy cadmodeling for assurance and confidence. Finishedmanufacturing includes compressed injection molding with a polycarbonate plastic material ensuring reliability (Fig. 44.4). Parts 1. Basket\u2014A self-adjusting designer bottom to accommodate a range of drinking glasses and cups. Bi-laterally placed slit is to settle up the handle of any type of cup. The main basket is with height 57 mm developed with reference to the smallest glass that we considered in our range, which is a paper teacup with a height of 64 mm. Thus, even the shortest cup is significantly accessible. Importantly, the pivot point should be higher than the center of gravity in order to maintain equilibrium"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003041_s10845-021-01803-1-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003041_s10845-021-01803-1-Figure6-1.png",
        "caption": "Fig. 6 a Layout of the multi-robot optical inspection system for leftside auto body; b Layout of MP locations on the auto body",
        "texts": [
            " To evaluate the effectiveness of the proposed method, we conducted a case study in an inline multi-robot inspection station for auto bodies. In the inspection cell, FANUC 2000iB-125L industrial robots are mounted with optical probes for dimensional deviation inspection. Four robots are denoted by R1, R2, R3 and R4, respectively. The locations of these sensing robots for the multi-robot optical inspection station are shown in Fig.\u00a02. The layout of the optical sensing station for the left-side auto body, and the locations and surface normal directions of MPs are shown in Fig.\u00a06. In Fig.\u00a06b, 45 MPs with different surface orientations, numbered from M1 to M45 are shown. The proposed method has 1 3 been programmed to allocate the MPs, optimize the independent path of each robot, and finally modify the path to get an optimal collision-free coordination path. The parameters used in the case study are listed in Table\u00a01. The fixed priority method and the delay startup method were introduced for comparison analysis in multi-robot coordination (Chen & Li, 2017; Qadi & Goddard, 2005; Shin & Zheng, 1992)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001845_s00202-020-01132-1-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001845_s00202-020-01132-1-Figure3-1.png",
        "caption": "Fig. 3 Different locations of the air-cooling speed measurement",
        "texts": [
            " This tool can measure the rates in the range of 0 to 30 m/s, and the dimension of probe diameter of the tip is 7 mm, which is selected according to the average fin space of TEFC housing. This type of velocity meter has an option called flow setup to choose the appropriate duct shape, e.g., round duct, rectangular duct, and so on. Furthermore, it has another option as actual and standard setup that the user can also select the standard temperature, standard pressure, and a source for the actual temperature. By using these options, the air velocity meter is calibrated. According to Fig. 3, to evaluate the mean speed of an aircooling flow in the circumference of the machine housing, the airflow speed was measured in three different positions along the axial direction of the cooling fins: at the beginning, in the middle, and at the end of the fins. To increase the measurement accuracy of the mean value of velocity, the velocity samplings are more than 60 points for each fan rotational speed. Figure 4 illustrates the air speed distributions along the axial direction of the several semi-open fin channels for the machine under study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000326_012006-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000326_012006-Figure1-1.png",
        "caption": "Figure 1. 2DOF model of the rolling bearing.",
        "texts": [
            " The balls which enter the loaded zone are deformed and produce varying flexibility vibration. The rolling element bearing is the most critical element of rotating machines, therefore its proper description of the bearing is of high importance from the viewpoint of the dynamic behaviour of the entire machine. The key components of any vibratory system are the mass, stiffness, damping and the external forces. Therefore, the deep groove ball bearing with a single row of balls is simplified and modelled as a nonlinear spring-damper system (figure 1). In the model, the outer ring was fixed in a rigid, non-deformable housing (the model considers the case of a non-rotating outer race), and the mass of the balls is neglected. The inner ring rotates together with the shaft. The elastic deformation between the raceways and each point of contact with the ball is assumed to be Hertzian. The assumptions and simplifications of the model are: \u2022 the ball bearing model has equi-spaced balls rolling on the surface of the inner and outer raceway and there is no interaction between them, \u2022 the slipping effect of the balls is neglected, \u2022 the motion of race and balls occur in the bearing\u2019s plane, \u2022 the deformations at the contacts are described by the Hertzian contact theory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000189_ecce.2019.8912698-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000189_ecce.2019.8912698-Figure3-1.png",
        "caption": "Fig. 3. Velocity field contours inside the slotted air gap at Ta=31166",
        "texts": [
            " Software Validation As a validation of the model, the results obtained numerically are compared to those calculated from empirical correlations of the literature. The difference between numerical and analytical results is reasonable and leads to accept that CFD method can be used to estimate windage losses. IV. ANALYSIS AND RESULTS A. Velocity field The rotor slot leads the airflow to generate vortices within the slot. The airflow is assumed to have the same velocity of the boundary at fluid\u2013solid boundaries. The flow is rotating at the same velocity near the rotor and is fixed near the stator. In Fig. 3 and Fig. 4 the velocity field is represented, from the stationary frame, for the rotating configuration at \u2126=20000 rpm (Ta=31166). The airflow is initially regular in the smooth part. At the slot, the change in thickness of the air gap makes the streamlines deviate towards the bottom of the slot with a swirling motion. A stagnation zone with a weak airflow recirculation can be detected and Taylor-Couette structures can as well be observed inside the slot. As the rotation speed increases, the turbulence level inside the slot increases accordingly",
            " 2, the skin friction coefficient Cf is expressed as a function of the wall shear stress i which is calculated by the CFD model. The coefficient can be calculated for each subsurface of the rotor from S1 to S5. As shown in Fig. 6 the friction coefficients of each slot sub-surface Cfk (k from 1 to 5) increase with the rotation speed of the rotor. It also shows that the most important friction coefficient appears on the smooth parts of the air gap S5 and S1 out of the slot. Cfk are less important at the bottom of the slot on S3. This is in agreement with the airflow velocity distribution shown in Fig. 3 revealing that surfaces on the low part of the slot are less exposed to the turbulent flow. D. Estimation of windage losses Windage losses can be estimated as defined in Eq. 1. for each sub-surface of the rotor. The sum of these losses on all rotating surfaces leads to the windage losses in the air gap of the electrical machine. Fig. 7 represents a comparison between windage losses calculated from analytical correlations and numerical results obtained by the CFD-LES method. Relatively few difference between results, demonstrates the ability of CFD methods to analyze the windage losses inside air gaps of electrical machines"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002244_012013-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002244_012013-Figure2-1.png",
        "caption": "Figure 2. Schematic drawing of test rig.",
        "texts": [],
        "surrounding_texts": [
            "Experimental studies are aimed at measuring the effect of both surface roughness and axial distance between the moving disk and the housing on the loss of friction power of the disk in centrifugal pumps. Most experiments in this area are based on the work of Pantell [2]. The test rig consists of one or two bearings 1 mounted on a shaft to which the lens or cylinder 2 to be tested is mounted. The outer part is a housing 3 consisting of one or more cylindrical parts centered relative to each other. The side windows are pressed against the ends of the cylindrical part with anchor bolts. Since the moment of friction of the fluid is measured in the housing, the moving parts must be able to rotate around the axis as smoothly as possible without contact with the stator parts. According to Fottinger's instructions, this is achieved by attaching stuffing box seals in a special housing. The housing 3 is attached to the frame 4 by means of anchor bolts 5. The inlet allows you to supply fluid close to the axis. At the highest point of the case there is a ventilation valve with a funnel, at the lowest point - a drain valve. The test rig is a closed loop. The pump inlet is connected to the tank via a flexible connection. The pump and the test disk are connected to the motor 6 through the clutch 7 torque meter 8, which facilitates the measurement of transmitted torque, and a special speed sensor for measuring the speed 9 [3]. HERVICON+PUMPS 2020 Journal of Physics: Conference Series 1741 (2021) 012013 IOP Publishing doi:10.1088/1742-6596/1741/1/012013 The aim of the research is to determine the influence of surface roughness and axial clearance on disk losses in the pump. The surface roughness of the rotating disk is changed by machining or by gluing sandpaper with different surface roughness. The axial clearance is changed by reducing the thickness of the disk by machining. It is proposed to use a Taylor-Hobson device with a diamond pin to measure the surface roughness of the impeller, hard disk and sandpaper used. Disk friction power losses are defined as the difference between the measurement results when the housing is filled with water and the power measured when the housing was empty. So energy losses in the bearings will be eliminated from the net power loss. As suggested by Bennett and Worster [4], Worster [5], and Varley [6], some of the energy expended on disk friction eventually returns to the impeller and helps combat bulk losses. To facilitate experimental research, the fraction of disk friction power added to the impeller power can be considered equal to the volume loss [7]. Against the general background, the work of Hergt and Prager stands out [8]. They measured the friction losses of the disk in real pumping conditions. An experimental installation with a bifurcated disk was taken. Thus, the front and rear walls of the cover were driven by two different engines running at the same speed. Thus, the power of the disk friction absorbed by the rear wall could be measured separately from the power of the impeller. The conditions of supply to the lateral sinus were determined by the speed of exit from the impeller [9]. This approach is necessary because the fluid flow conditions are completely different from those shown in the tests. The main flow at the outlet of the channels of the impeller has a large circumferential velocity and any flow into the lateral sinus has a momentum. It seeks to accelerate lateral flow and reduce disk friction."
        ]
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure23-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure23-1.png",
        "caption": "Fig. 23 Stress analysis of the upper panel",
        "texts": [
            " The softwarebased optimised design (SBDOP) shows a 41\u00a0MPa increase in stress at the left-hand side body attachment compared to the baseline. The EBDOP-1 and EBDOP-2 designs show an ~ 350\u00a0MPa decrease at the attachment point but show additional local stress hotspots above the yield stress (350\u00a0MPa). The software-based optimised design (SBDOP) shows a 129\u00a0 MPa increase in stress at the right-hand side body attachment compared to the baseline. The EBDOP-1 and EBDOP-2 designs show an ~ 350\u00a0MPa decrease relative to the baseline at the attachment point. The stress analysis of the upper panel is shown in Fig.\u00a023 and presented in Table\u00a05. Expert-based optimisation-1 (EBOP-1) is better method to reduce stress for upper panel. 7.1.4.1 Von\u2011Mises Stress Overview All elements in red are above the assumed yield stress of 350\u00a0MPa. The softwarebased optimised design (SBDOP) shows a 290\u00a0MPa increase in stress at the left-hand side body attachment compared to baseline but a reduction in the number of hotpots above 350\u00a0MPa. The EBDOP-1 and EBDOP-2 designs show an 1 3 ~ 890\u00a0MPa increase at the attachment point, with additional hotpots and a larger area of the lower panel with stresses > 350\u00a0MPa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002037_j.engfailanal.2020.105126-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002037_j.engfailanal.2020.105126-Figure17-1.png",
        "caption": "Fig. 17. Upper arm of the tower (arm 1 and 2).",
        "texts": [
            " If the input is taken into account in the Table 6, then, the value of X, is the thickness of ice X for a coefficient of variation of 20% and a probability of 0.2% to be exceeded in a year or a return period of 500 years using the following Eq. (8) and Table 7. X = X\u2019 \u2212 0.45\u03c3 + \u03c3 1.282 [ \u2212 ln( \u2212 ln(1 \u2212 p(X)) ) ] (8) The designs were made taking into account stress, dead and wind loads for suspension structures (Fig. 8 and Fig. 9) are the Fig. 14, Fig. 15, Fig. 16 and the retention (Fig. 11 and Fig. 12) structures are in the Fig. 17, Fig. 18. The wind load was estimated following the procedure established by this paper in the section 2 and Eq. (9). Load = 0.613 \u00d7 V2Vkz \u00d7 GRF \u00d7 I \u00d7 Dd \u00d7 A (9) Where: Fig. 11. Retention structure for normal condition. Fig. 12. Retention structure for construction condition. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 0.613: It is the constant for metric unit for the density in the air. It reflects the air density for 1 atmosphere, 15 C, with mercury pressure 760 mm, in the Table 8",
            " 21; besides, the structural model of the TO type towers generated in the Fig. 22, Fig. 23, Fig. 24, Fig. 25, Fig. 26, Fig. 27 and Fig. 28. In the Fig. 14, the suspension tower is modeled with finite element; each part of the tower has been built for the validation of the loads according to the section 3. In the Fig. 15, the material for retention tower has built, with the material influence for the calculations and constraints as elastic module, yield stress, steel type and components. Each component of the tower has been described in the Fig. 16 with the tower top. Fig. 17 for both upper arms, it applies to arm 1 and 2, Fig. 18 with the lowest arm of the tower, Fig. 19 in the upper body between arm 1 and 2, Fig. 20 for the body structure between arm 2 and 3, Fig. 21 and Fig. 22. Fig. 14. Tower model in finite element. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 The stage division of the ice disaster is composed of four stages: Preconditions (associated to the pre-failure period), steady-state progression and multiple failures (during disaster period) and restoration (post-failure period) [38]; recent papers have indicated effective resilience enhancement framework for towers installed, however, the recommendations for new assets are not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure11-1.png",
        "caption": "Fig. 11. Open-circuit field distribution of the CPM-II with (a) \u03b4=15 deg, (b) \u03b4=30 deg, (c) \u03b4=45 deg, (d) \u03b4=60 deg, (e) \u03b4=75 deg, (f) \u03b4=90 deg, (g) \u03b4=105 deg, and (h) \u03b4=120 deg.",
        "texts": [],
        "surrounding_texts": [
            "1) Air-gap flux density Fig. 9 shows the variation of fundamental airgap flux density of CPM-II with slotless stator. The airgap flux density first increases and then decreases as \u03b1 increases. The optimal \u03b1 is about 0.7, which is larger than that of CPM-I. In addition, the variation tendency of flux density with \u03b4 is similar to that with \u03b1, and the optimal \u03b4 is about 60 deg. Fig. 10 shows the fundamental air-gap flux density variation of CPM-II with variable \u03b4. It can be observed that when \u03b4=60 deg, the maximum fundamental airgap flux density can be achieved, meaning that the slotting effect cannot affect the variation tendency of flux density. 2) Magnetic Saturation The open-circuit and on-load field distributions of the CPM-II are shown in Figs. 11 and 12, respectively. It can be observed that the flux density in Region-II is significantly reduced and the saturation of the salient iron pole is weakened as \u03b4 increases. Hence, an excellent diffluence effect can be achieved, and the above theoretical analysis is verified. Figs. 13 and 14 show the on-load magnetic field distribution of CPM-I and CPM-II with \u03b4=90deg under different currents. It can be noticed that saturation of salient core of CPM-I is significant. However, the saturation of salient core of CPM-II with \u03b4=90deg is almost eliminated. (a) \u03b4 Fh Fhy Fhx 0 0 Fhx Fhy r r \u03b4 \u03b4\u03b4 lw Fhx-max Fhy-max PM1 PM2PM2 Stator Air-gap \u03b41 \u03b42 Leakage flux \u03b41 \u03b42 Authorized licensed use limited to: Rutgers University. Downloaded on May 20,2021 at 12:02:13 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (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. (c) \u03b4=45 deg, (d) \u03b4=60 deg, (e) \u03b4=75 deg, (f) \u03b4=90 deg, (g) \u03b4=105 deg, and (h) \u03b4=120 deg."
        ]
    },
    {
        "image_filename": "designv11_63_0002504_j.ijmecsci.2021.106392-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002504_j.ijmecsci.2021.106392-Figure4-1.png",
        "caption": "Fig. 4. Mechanical model of the hoop.",
        "texts": [
            " Due to he invariance of the system to translations and rotations on the plane, ny straight trajectory with the heavier mass placed ahead corresponds o a stable reference motion. In this analysis, the forward velocity has een considered positive, directed from mass \ud835\udc34 to \ud835\udc35. .2. Hoop rolling without slipping Secondly, the linear stability of the forward upright motion with contant velocity of a hoop rolling without slipping is addressed resorting o the procedures of Section 2 . .2.1. Mechanical model and equations of motion of the hoop The hoop can be modelled with five coordinates: \ud835\udc65 \ud835\udc36 and \ud835\udc66 \ud835\udc36 locate he contact point, denoted by \ud835\udc36 in Fig. 4 ; \ud835\udf13 is the yaw angle; \ud835\udf19 represents he lean angle, and \ud835\udf03 corresponds to the roll angle. Therefore, the set of eneralized coordinates is: = ( \ud835\udc65 \ud835\udc36 \ud835\udc66 \ud835\udc36 \ud835\udf13 \ud835\udf19 \ud835\udf03 )T . (69) responds to body 2. The orientation of the intermediate frames \u27e8\ud835\udc4b \u2032\ud835\udc4c \u2032\ud835\udc4d \u2032\u27e9, \u27e8 ned by the matrices \ud835\udc79 \u2032, \ud835\udc79 \u2032\u2032 and \ud835\udc79 2 , which can be expressed in terms of the e \ud835\udc79 (70) T ass \ud835\udc3a are given by: \ud835\udc93 (71) w mic constraints. The velocity of the contact point can be computed as: \ud835\udc97 (72) w , the nonholonomic constraints are: \ud835\udc6a (73) N o holonomic constraints, and hence the number of degrees of freedom is \ud835\udc5b nlinear index-2 DAE system which, by differentiating once with respect to t E system: ( (74) The inertial frame is designated as body 1, whereas the hoop cor \ud835\udc4b \u2032\u2032\ud835\udc4c \u2032\u2032\ud835\udc4d \u2032\u2032\u27e9 and the body frame 2 \u27e8\ud835\udc4b 2 \ud835\udc4c 2 \ud835\udc4d 2 \u27e9, shown in Fig. 4 , is determi lemental rotation matrices as follows: \u2032( \ud835\udc99 ) = \ud835\udc79 \ud835\udf13 , \ud835\udc79 \u2032\u2032( \ud835\udc99 ) = \ud835\udc79 \ud835\udf13 \ud835\udc79 \ud835\udf19, \ud835\udc79 2 ( \ud835\udc99 ) = \ud835\udc79 \ud835\udf13 \ud835\udc79 \ud835\udf19\ud835\udc79 \ud835\udf03 . he absolute position vectors of the contact point \ud835\udc36 and the centre of m \ud835\udc36 = \u239b \u239c \u239c \u239d \ud835\udc65 \ud835\udc36 \ud835\udc66 \ud835\udc36 0 \u239e \u239f \u239f \u23a0 , \ud835\udc93 \ud835\udc3a = \ud835\udc93 \ud835\udc36 + \ud835\udc79 \u2032\u2032 \u239b \u239c \u239c \u239d 0 0 \ud835\udc45 \u239e \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239d \ud835\udc65 \ud835\udc36 + \ud835\udc45 sin ( \ud835\udf13 ) sin ( \ud835\udf19) \ud835\udc66 \ud835\udc36 \u2212 \ud835\udc45 cos ( \ud835\udf13 ) sin ( \ud835\udf19) \ud835\udc45 cos ( \ud835\udf19) \u239e \u239f \u239f \u23a0 , here \ud835\udc45 is the radius of the hoop. The assumption of rolling without slipping results in two nonholono \ud835\udc36 = \ud835\udc97 \ud835\udc3a + \ud835\udf4e 21 \u00d7 \ud835\udc93 \ud835\udc3a\ud835\udc36 \u21d2 \ud835\udc97 \ud835\udc36 = \u239b \u239c \u239c \u239c \u239d \ud835\udc63 \ud835\udc36 \ud835\udc65 \ud835\udc63 \ud835\udc36 \ud835\udc66 \ud835\udc63 \ud835\udc36 \ud835\udc67 \u239e \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239d "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001471_tase.2020.3015930-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001471_tase.2020.3015930-Figure1-1.png",
        "caption": "Fig. 1. Prototype of robotic transtibial prosthesis. (a) Robotic prosthesis and (b) mechanical structure of prosthesis. The components of the prosthesis include control circuit, sensors (one strain gauge, one angle sensor, and one IMU), battery, carbon-fiber foot, dc motor, gear box, and ball screw.",
        "texts": [
            " [19] have developed an adaptive algorithm by designing a filter by off-line analysis and updating threshold using newest gait data to detect the gait events (initial contact, end contact, and mid swing). However, their study still requires some off-line operations to get some parameters, which will bring other device (i.e., computer) and need human interference. Besides, its adaptation for different people needs further validation. Here, we utilized a pattern recognition method, instead of threshold decision method, and designed an adaptive method to detect MDF timing. 1) Prosthesis Prototype: As shown in Fig. 1, a commercialized robotic transtibial prosthesis produced by SpeedSmart Company Ltd (a spin-off company of Peking University) was used in this study. The prosthesis is comprised of a control circuit, sensors [including one full bridge of strain gauge, one angle sensor, and one inertial measurement unit (IMU)] and battery, and its weight is about 2 kg. More details can be found in [2], [9], and [16]. The full bridge of strain gauge of prosthesis is used to record the deformation of carbon-fiber foot, which can be used to detect the stance and swing phases in each gait cycle [20]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure1-1.png",
        "caption": "Fig. 1. Preliminary design of conductive traces which mate to a standard 9V battery (red for the positive terminal and dark grey for the negative).",
        "texts": [
            " The latter doubles as a means of maintaining the battery firmly within the mechatronic assembly. To this end, a snap interference fit is employed, as it ensures contact without the use of extra components [18]. In order to simplify the printing process, the conductive and non-conductive materials can alternatively be printed separately, opening up for single extruder printer use. The conductive parts should mate with the battery terminals, taking into account printer-specific tolerances, as revealed in Figure 1). The 3D printer used in this study was a Flashforge Dreamer [19] hobby printer, with manufacturer-claimed tolerances of \u00b10.2 mm in-plane and \u00b10.5mm perpendicular to the build plate. The matching dimensions that would ensure a reasonable fit between battery terminals and conductive traces were obtained by trial and error, as batteries are produced with too large a geometrical tolerance range to enable accurate predictive design [20]. As a metric of this uncertainty, the positive terminal outer diameter is measured to be 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002396_0142331220987532-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002396_0142331220987532-Figure4-1.png",
        "caption": "Figure 4. Configuration of the ME type sensor positioning structure in the magnetic field zone in the machine.",
        "texts": [
            " The equation of the voltage generated in open circuit: V0 t\u00f0 \u00de= laV B t\u00f0 \u00de m0 \u00f025\u00de The equation of the load voltage V= laV H3RL 1 jwC +RL = laV B3RL WC m0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1+ wCRL\u00f0 \u00de2 q \u00f026\u00de The equation of the optimal resistance of the electric circuit Ropt = 1 wC = tC 1 n\u00f0 \u00de nSE 11 + 1 n\u00f0 \u00deSH 33 2pfwle33 n 1 K2 31 SE 11 + 1 n\u00f0 \u00deSH 33 \u00f027\u00de The electromechanical conversion of energy is located in the gap. This conversion is therefore affected by any magnetic imbalance, mechanical, electrical or electromagnetic rotor or the stator. The flux hugged in the stator windings or the leakage flux in the axis of the rotor are parameters that, because of their sensitivity to any imbalance of the machine, deserve to be analyzed (Figure 4). If we consider Ncc the number of turns in short circuit and icc (t) the instantaneous current in these, then the pulsating magnetomotive force Fcc associated with a winding having short-circuited turns can be expressed by Fcc u, t\u00f0 \u00de= 1 p NccIcc X\u2018 n= 1 1 n sin np 2p cos vs t u6n u\u00f0 \u00de \u00f028\u00de With: icc(t)=Icc cos(vs t u) This brings us to the expression, in the reference linked to the rotor, of the induction of gap to identify the frequencies of interest for the detection of short circuit from the axial flux Ber, cc = X\u2018 n= 1 X\u2018 k = 1 Bncc cos k6n 1 g P vst u6nu0 \u00f029\u00de Bncc is the amplitude of the ni eme harmonic, k an integer associated with the harmonics of the alimentation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure5-1.png",
        "caption": "Fig. 5. Principle of COG measurement.",
        "texts": [
            " (2) Calculate the position of pendulum axis in RCS: since coordinates transformation matrixes between BCS and PCS are obtained, the position of pendulum axis in RCS can be calculated using coordinates transformation matrixes between RCS and BCS obtained in Step 1st [8,24]. Step 3rd: Measure oscillation period of the measured body. Make the measured body swing around the pendulum axis in a small angle and measure the oscillation period of the measured body. Step 4th: Repeat the measurement and process the measured data. (1) Change suspended postures of the measured body and repeat Step 2nd to Step 3rd for 10 to 12 times. (2) After finishing the measurement, several pendulum axes in RCS are obtained as shown in Fig. 5. L1, L2 and L3 in Fig. 5 are three pendulum axes in RCS Or-XrYrZr under three different suspended postures of the measured body respectively. The COG of measured body falls on pendulum axis automatically because of the effect of universal joint, so in theory all pendulum axes will intersect at the COG. However, for the reasons of measurement errors, COG is determined using methods presented in Section 2.4. (3) Calculate inertia tensor of the measured body defined in CCS using Eq. (3), (5a) and (6). The inertia parameters of a measured body are obtained after finishing the procedures described above"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002521_j.euromechsol.2021.104267-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002521_j.euromechsol.2021.104267-Figure3-1.png",
        "caption": "Fig. 3. Phenomenological mechanism and dynamic model with 1DOF.",
        "texts": [
            " (13) and (14) depend only on the relative angles and \u03bbi, being crucial for the shape of the nonlinear torque curves that will be presented in Section 5. Function Gi introduces the torque changes due to the orientation of the static load, and Di also takes into account the effects related to the deformation of the spring. The mechanism from Yang et al. (2020) is described by similar equations. Thus, different torque curves can be obtained based on the combination of springs: p1T = \u2211n i=1 p1i (15) p2T = \u2211n i=1 p2i (16) A phenomenological mechanism is represented in Fig. 3a, where input torque p2 is applied to the outer track side. For this work, the connections at both tracks allow the springs to rotate without any resistance. Fig. 3b shows the nonlinear link attached to a 1 DOF system, whose inertia, torsional stiffness and damper coefficients are described by J, kt, and ct, respectively. Without the nonlinear link, its natural frequency and damping factors are calculated as \u03c91 = \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 kt/J \u221a and \u03b6 = c/(2J\u03c91), H.H. Miyasato et al. European Journal of Mechanics / A Solids 89 (2021) 104267 respectively. Despite its simplicity, this minimal model was designed to provide a comprehensive insight about the interaction between the mechanism and a pre-existing system. Assuming that the inertia is connected to the inner track (Fig. 3a), the equations of motion take into account the contribution of the output from Eq. (15): J\u03b8\u03081 + ct\u03b8\u03071 + kt\u03b81 = \u2211n i=1 p1i (17) Finally, the input is assumed as oscillatory at frequency \u03c9: \u03b82(t)=\u03982cos(\u03c9t) (18) Considering a normalization according to time (\u03c4 = \u03c91t), the system can be rewritten in terms of non-dimensional coefficients: d2\u03b81 d\u03c42 + 2\u03b6 d\u03b81 d\u03c4 + \u03b81 = p(\u03c4)= \u2211n i=1 ( \u03c12 i Di + uiGi ) (19) , where \u03c1i = \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 kiR2 2i kt \u221a and ui = 0fki R2i \u03c92 1J are named as input and static ratios, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001558_s10778-020-01007-9-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001558_s10778-020-01007-9-Figure3-1.png",
        "caption": "Fig. 3 Fig. 4",
        "texts": [
            " It can be seen that the study of the evolution of the system in time was carried out by sequential numerical integration of the system of equations (1.2) at different stages of motion. In this case, the final conditions of the current stage were the initial ones for the next one. The initial conditions for the motion of the system at the first stage were chosen to be zero: in the calculations, we studied the system motion at rest. 2. Analysis of the Results of Modeling. We obtained graphs that describe: the speed of the crane (Fig. 2), the displacement of the load on the flexible suspension from the vertical (Fig. 3), the optimal driving forces implemented by the frequency-controlled electric drive (Fig. 4). The graphs that correspond to the optimal control of the crane with classical control constraints (expressions (2.6) in [13]) are shown in gray, and those with modified constraints (expressions (2.7) in [13]) are shown in black. To compare the results of modeling the optimal controls for a crane with a load on a flexible suspension, the characteristics of the overload cycle are collected in Table. 2.1. All the characteristics were obtained using the parameters of the harbor bucket unloader given in Table 3",
            " However, the function \u00d2 t f u 0 ( max , u min ) of the arguments u max and u min that are set taking into account the possibility of practical implementation of the control remains unknown. Considering the practical importance of time-optimal control problems for dynamic systems, determining the influence of constraints on the speed control of systems is an urgent scientific problem that requires further theoretical study. Figure 4 shows that the function of the optimal driving force that corresponds to the optimal control of the crane with a load has a high-frequency component. It results from the free currents of the electric drive of the crane motion mechanism. However, Fig. 3 shows that the effect of electromagnetic processes on the oscillations of the load is not significant; therefore, in such problems, they can be neglected. The transition from classical tomodified control constraints allows reducing the average value of the drive force by 1.50 times in acceleration mode and by 2.37 times in deceleration mode. This leads to a decrease in the dynamic loads in the mechanical transmissions of the cranemotionmechanism and in the cranemetal structure and increases the crane\u2019s durability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure7-1.png",
        "caption": "Fig. 7. Installation diagram of the middle phase device",
        "texts": [
            " Looking at the structure at the middle phase position from the right side of Fig.6, you will find two V-shaped angles protruding outwards. When the surface of the right device is a complete mesh surface, it will be blocked by the angle steel, and cannot fit with the vertical surface of the angle steel where the hanging point is located, and there must be a gap on the right side. We have already introduced that middle-phase\u2019s anti-bird device is composed of two parts, both of which have a triangular prism shape. The state of the device after installation is shown in the Fig.7. According to the modeling diagram of the middle phase part of the pole tower, we can find that the right side of the middle phase has an impact on the sealing of the device due to the presence of the angle steel structure, which is the cause of the gaps in the existing antibird devices. Which may cause small birds to enter the middle phase protection area through the gap. The left side does not have the influence of the angle steel, and the ordinary triangular prism device can protect the left side area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002290_jestpe.2021.3057718-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002290_jestpe.2021.3057718-Figure1-1.png",
        "caption": "Fig. 1. Phasor diagram of the SPMSM in steady state. (a) Based on rotor-fieldorientation. (b) Based on square-wave angle injection under id1 = 0 control.",
        "texts": [
            " The measured data including d-, q-axes currents, voltages, and electrical angular speed are filtered by second-order butterworth lowpass filters respectively. Online parameter estimation could be done based on the steady state machine model after filtering [12], which could be simplified as follows: ( ) ( ) ( ) ( ) ( )q q s d mu k Ri k L k i k k  = + + (2a) ( ) ( ) ( ) ( )d d s qu k Ri k L k i k= \u2212 (2b) where k is the index of the discrete sampling instants in the estimation system. As is shown in Fig. 1(b), \u2206\u03b8 is defined be positive when square-wave angle is injected. It would lead rotor-field-orientation in dq1 reference frame. Equation (2) is represented as ( ) ( ) ( ) ( ) ( )cos( )q q s d mu k Ri k L k i k k   = + +  (3a) ( ) ( ) ( ) ( ) ( )sin( )d d s q mu k Ri k L k i k k   = \u2212 +  (3b) where \u2206\u03b8 is the square-wave angle. Rotor flux linkage and stator inductance are parameters to be estimated in (3). The other variables, including electrical angular speed, d-, q-axes currents and voltages, are measured",
            " As is shown in [29] and [30], the average voltage errors of d- and q-axes are expressed as _ dead _ d err d q err q u D V u D     =        (4) where Dd and Dq are d- and q-axes distorted coefficients respectively, Vdead represents the magnitude of the voltage error, which is defined as dead on off sat d dead dc sat d( ) 2s T T T V V V V V V T + \u2212 + = \u2212 + + (5) where Tdead is dead-time, Ton and Toff are turn on and turn off time of the switching device, Ts is the switching period, Vdc is measured real-time DC bus voltages, Vsat and Vd are the forward voltage drop of the switching device and the diode. Dd and Dq are represented as 1 sgn( ) ( ) sgn( ) ( )sgn ( ) sgn( ) a d d e b e e q q c i D i i D i i   \u2212          = =               K K K (6) where 2 2 sin sin( ) sin( ) 2 3 3 ( ) 2 23 cos cos( ) cos( ) 3 3 e e e e e e e              \u2212 +  =    \u2212 +    K , K-1(\u03b8e) is the inverse of K(\u03b8e), \u03b8e is electrical position, Dd and Dq are the function of rotor position and \u03b3, where \u03b3 represents the angle between the stator current vector and q-axis in the synchronous rotating reference frame as shown in Fig. 1(a). After square-wave angle \u2206\u03b8 is injected, Dd1 and Dq1 are represented as Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 17,2021 at 00:05:22 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (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. 1 1 sgn( ) ( ) sgn( ) sgn( ) a d e b q c i D i D i         = +          K (7) According to the equation (6) and (7), the Ddq and Ddq1 are functions of rotor position and \u03b3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000431_ab6158-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000431_ab6158-Figure1-1.png",
        "caption": "Figure 1. First prototype of the stand-alone device (a) and particular of the board (b).",
        "texts": [
            " In section 2 the features of the stand-alone device are presented, while in section 3 the mathematical model of the system is derived and the LQR and Kalman filter are defined. In section 4 the state recovery algorithm is explained and the synthesis of the controller is made. In section 5 the configuration of the system which has been used to perform the numerical analysis is presented and the corresponding results are shown. Finally, in section 6 the experimental results obtained on a large vibrating system are shown, while conclusions are drawn in section 7. The developed device (shown in figure 1) is able to carry out operations of vibration control in an autonomous way [16, 17]. Two MEMS digital accelerometer are available: one is placed on the fixed frame of the device and the other on the suspended mass of the actuator. The first sensor measures the acceleration of the structure, which will be used to compute the control signal, while the second one measures the acceleration of the actuator, in order to get information about its dynamics during the control action. These signals are acquired by the microprocessor, the main element of the electronic board in which different control algorithms can be implemented [18\u201320]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure14-1.png",
        "caption": "Figure 14. Third and fourth eigenmode (95.8 Hz and 129.45 Hz).",
        "texts": [],
        "surrounding_texts": [
            "Composite materials, finite element analysis, automotive engineering, vibrational behavior, design optimization Date received: 21 March 2021; accepted: 5 May 2021"
        ]
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure2-1.png",
        "caption": "Fig. 2. Two types on contact point after ring displacement.",
        "texts": [
            " Jose established a 3-DOF statics model for 4PBB [11], and Halpin established a quasi-static analytical model for 4PBB [12]. At present, although the statics model of the ball bearing is very mature, the description of the geometric relationship of steel ball and raceway is too brief. Usually, the deformation of each steel ball at each azimuth position is directly obtained from the curve equation of the raceway curvature center locus [7, 8]. However, according to the displacement of the steel ball, the contact deformation caused by the radial displacement has at least two types, as shown in Fig. 2. Fig. 2 shows an example of the displacement of ball-raceway contact points, while the outer ring has pure radial movement. Outer ring contact point type 1 means the steel ball is sliding on the outer raceway surface; outer ring contact point type 2 means the steel ball is stuck to the original contact point with the outer raceway. The difference between these two assumptions is usually relatively small, but in the case of large radial displacements, the modeler needs to have a full understanding of these differences"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002629_012013-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002629_012013-Figure6-1.png",
        "caption": "Figure 6. Motor housing to end cover fan",
        "texts": [],
        "surrounding_texts": [
            "Modal Analysis is mostly dispensed to seek out natural frequency of the system and completely different mode shapes at different Eigen values [12,14]. In this work, totally different modes and Eigen values square measure calculates and premeditated between mode shapes versus natural frequencies [15,20, Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 21]. This shows the stiffness and rigidity of the material toward totally different modes. Using Modal analysis it can be captured the first 10 natural frequencies and related mode shape of each frequency for the 3 cases. Figure 7 shows the modal analysis results for case A. Each figure represents the frequency of each mode and displacement related to the frequency. Due to rotational velocity the model is exhibiting rotational vibration. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Table 1 shows the value of modal analysis for case A, B and C. According to this table Case C is having more natural frequency than other 2 cases. Figure 8 shows the relation between modes and damped frequency. Based on modal analysis results Case C (Hybrid Hemp - Flax fibre) has additional stiffness (due to high natural frequency) while compared to other two cases. Conclusion In this analysis, a study for three totally different composite materials particularly Hemp fibre, Flax fibre and Hybrid of Hemp and Flax fibre for Finite Element Analysis of electrical motor casing has been carried out. The motor casing should be rigid and better in stiffness to resist operative vibrations. 3D modelling of motor casing is built using Parametric Creo one among the fine 3D modelling software. These 3 materials are applied to motor casing as a material properties of the model and three types of analysis namely, Modal Analysis, Rotor Dynamics and Structural Transient Analysis carried out to check the stiffness, reliability towards operative frequency and durability of motor casing. As determined from modal analysis Case C (Hybrid Hemp Flax fibre) has better stiffness compared to two different fibres as shown in graph fig 7 (Modes vs Damped Frequency). Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Campbell diagram was premeditated between frequencies vs speed of the motor for predicting critical speed. The operative frequency and operative speed do not match with the critical speed and frequency. Model analysis was performed for finding the vibrational stability of the composite materials and observation Case C shows low deflection and equivalent stresses compared to different fibres. Force vs deflection curve and stress vs strain curve are premeditated to indicate the stiffness of the materials and energy vs deflection curve premeditate show of the strain energy observed, when deformation happens. As determined from all above studies Case C has higher stiffness and strength compare from different two fibres. Reference [1] Bajuri F, Norkhairunnisa Mazlan, Mohamad Ridzwan Ishak and Junichiro Imatomi, 2016, \u201cFlexural and compressive properties of hybrid kenaf/silica nanoparticles in epoxy composite\u201d, Procedia Chemistry, Vol.19, pp:955-960. [2] Korniejenko. K, Frczek. E, Pytlak. E and Adamski. M, 2016, \u201cMechanical properties of geo polymer composites reinforced with natural fibres\u201d, Procedia Engineering, Vol.151, pp:388393. [3] Lut Pil, Farida Bensadoun, Julie Pariset and Ignaas Verpoest, 2016, \u201cWhy are designers fascinated by flax and hemp fibre composites?\u201d, Composites: Part A, Vol.83, pp:193-205. [4] Shuhimi F, Mohd Fadzli Bin Abdollah, Kalam M.A and Hilmi Amiruddin, 2016, \u201cTribological characteristics comparison for oil palm fibre/epoxy and kenaf fibre/epoxy composites under dry sliding conditions\u201d, Tribology International 101. [5] Raman Bharath, Vijaya Ramnath B and Manoaran N, 2015, \u201cKenaf fibre reinforced composites: A review\u201d, ARPN Journal of Engineering and Applied Sciences, Vol.10, No.13, pp:5483-5485. [6] Zamria M, Hazizan Md Akil and Zainal Ariffin Mohd Ishak, , 2016 \u201cPultruded kenaf fibre reinforced composites: Effect of different kenaf fibre yarn tex\u201d, Procedia Chemistry, Vol.19, pp:577-585. [7] Bensadoun. F, Depuydt. D, Baets. J, Verpoest. I, Van Vuure. A.W, 2017, \u201cLow velocity impact properties of flax composites\u201d, Composite Structures, Vol.176, pp:933-944. [8] Changlei Xia, Jason Yu, Sheldon Q. Shi, Ying Qiu, Liping Cai, Felix Wu. H, Han Ren, Xu Nie and Hualiang Zhang, 2017, \u201cNatural fibre and aluminum sheet hybrid composites for high electromagnetic interference shielding performance\u201d, Composites Part B, Vol.114, pp:121- 127. [9] Ming Liu, Andreas Baum, Jurgen Odermatt, Jens Berger, Liyun Yu, Birgitte Zeuner, Anders Thygesen, Jesper Holck and Anne S. Meyer, 2017, \u201cOxidation of lignin in hemp fibres by laccase: Effects on mechanical properties of hemp fibres and unidirectional fibre/epoxy composites\u201d, Composites: Part A, Vol.95, pp:377-387. [10] Bambach. M.R, 2018, \u201cGeometric optimisation and compression design of natural fibre composite structural channel sections\u201d, Composite Structures, Vol.185, pp:549-560. [11] Ravi Prakash Babu K., Rao P.K.V., Ali M.A., Raghu Kumar B., 2019, \u2018Crack simulations in shaft-hub-blade system using modal analysis\u2019, International Journal of Recent Technology and Engineering, 8(1), PP.2646-2650. [12] Bala Subramanyam P.N.V., Rao B.N., Prakash R.L., Rajavardhan D.S.K., Sankarudu B.S.K., Ganesh P.,2019, \u2018Modal analysis on go-kart chassis\u2019, International Journal of Engineering and Advanced Technology, 8(4), PP.1701-1705. [13] Syed K., Mannam H., Koya M.K., Meduri S.K., Laxmi V.S., Kasturi S.P., 2019, \u2018Dynamic simulation of gas turbine blade using finite element analysis\u2019, International Journal of Mechanical and Production Engineering Research and Development, 9(), PP.745-751. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 [14] Dama K.K., Surya S., Babu C., Pavan, Kiran R., 2018, \u2018Modal analysis of an automotive subsystem using cae techniques\u2019, International Journal of Mechanical Engineering and Technology, 9(3), PP. 1157-1162. [15] Deepak Kumar G., Kommuri N., Kota H., Prasanth V., Hameed S, 2018, \u2018Vibrational analysis of a human body on an automobile\u2019, Journal of Advanced Research in Dynamical and Control Systems ,10(6), PP. 39-45. [16] Devi S.K., 2018, \u2018The optimization of layer densities for multilayered insulation systems by using ansys\u2019, International Journal of Mechanical and Production Engineering Research and Development ,8(4), PP. 1087-1098. [17] Balasubramanyam P., Nageswara Rao B., Banerjee M., Lakshmana Swamy B., 2019, \u2018Impact analysis on Go-Kart chassis with variable speeds using ansys 19.0\u2019, International Journal of Engineering and Advanced Technology, 8(6), PP.2614-2620. [18] Kancheti N., Reddy Vemula A., Reddy Gudibandla G., Krishna H., Bala Subramanyam P.N.V. 2019, \u2018Modeling and analysis of wheel rim using ansys\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(8), PP.415-418. [19] Bharath E.S., NagaSree Harsha O., Bala Subramanyam P.N.V., Reddy G.S.S., 2019, \u2018Impact analysis on honey comb structured go-kart bumper using ansys R19.1\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(6), PP.396-400. [20] Ravi Prakash Babu K., Gupta S.K., Phaneendra Sai Sri D., Gautam K., Kumar C.M. 2017,\u2019Significance of modal analysis to detect and diagnose misalignment fault in turbine rotor\u2019,Journal of Advanced Research in Dynamical and Control Systems,9(Special issue 14),PP.2637-2648. [21] Babu K.R.P., Prasanth K.F.V., Sai V.V., Kumar M.V.S., Kumar B.N.S., Chaitanya B.S. 2017,\u2019Effect of pre-twist on free vibration characteristics of a cantilever beam with experimental validation\u2019,International Journal of Mechanical Engineering and Technology,8(5),PP:389-399. [22] T. Srinivasan, CH.V.L.N. Ram Gopal, P. Sasidhar, P. Vineeth Kumar and Shaik Sohail, 2020, Effect of change of cross section and Frequency Analysis of Anti-Roll Bar Using FEA, IOP Conf. Series: Materials Science and Engineering 954 (2020) 012027 IOP Publishing, doi:10.1088/1757-899X/954/1/012027"
        ]
    },
    {
        "image_filename": "designv11_63_0003002_s10846-021-01416-z-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003002_s10846-021-01416-z-Figure3-1.png",
        "caption": "Fig. 3 The principle of the lever motion according to commands signal when n is 8. [37]",
        "texts": [
            " The resolution of n-bit for the command signal is the default setting by selecting the integer n. The rest location of lever can be made by combining the signal without 2n\u22121. The range of 2n\u22121 + 1 to 2n \u2212 1 is used to locate the lever to be positioned to the front or to the right. The larger the value, the further the lever moves away from the origin. The range of 2n\u22121 \u2212 1 to 0 involves the same principle but acts in the reverse. The workspace of the manipulator within the predefined command signal is illustrated at the top of Fig. 1 when n is set by 8. And Fig. 3 briefly illustrates the principle of the lever motion according to commands, input signal and mechanism of the manipulator. The sub figure d) occurs when two commands signal is combined. The first three, four and bottom of Fig. 1 show the motion of the links of the excavator according to each direction of the lever. The further the lever moves away from the origin, the larger hydraulic force acting on the cylinder can be generated due to the more opened spool. In summary, the workspace of the lever is limited"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002175_978-981-15-7711-6_18-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002175_978-981-15-7711-6_18-Figure17-1.png",
        "caption": "Fig. 17 Z displacement plot at the end of forming",
        "texts": [
            " As the part design and the material are fixed and changes cannot be done, necessary changes are to be made to the forming process and the design of die considering the non-uniform springback. This forming process is suggested based on the FE analysis results. In Finite Element Analysis of Stamping Process \u2026 329 this process, dies are providedwith proper springback compensation.More than 10% thinning is not acceptable. Leading-edge and tip should lie on the same plane. The normal distance between the root plane and tip plane is 5 mm with a tolerance of \u00b11 mm. From Fig. 16 it can be said that maximum thinning is less than 0.5% which is negligible (well below the 10% limit) From Fig. 17, it can be observed that the portion of leading edge near root has deformed more compared to the portion of leading edge near tip. This is because of relatively greater spring back observed at that portion. As discussed previously, the spring back is not uniform and the compensation should be in accordance with the spring back. From Fig. 18 it can be seen that the leading edge and tip nearly lie on the same plane with acceptable deviation from the required. The portion at the centre of the leading edge seems to be deviating but the deviation is acceptable or can be sort out manually"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002113_0954407020984668-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002113_0954407020984668-Figure11-1.png",
        "caption": "Figure 11. Conceptual representation of the proposed reduced model.",
        "texts": [
            " For example, if it is assumed to use an interpolating seventh degree polynomial curve such as (15): f u\u00f0 \u00de= a1 u7 + a2 u5 + a3 u3 + a4 u \u00f015\u00de and it is assumed to consider Table 1 test case, the following values of the coefficients of equation (15) are obtained: a1 = 9:5524 10 16, a2 =7:0677 10 11, a3 = 1:5616 10 6, a4 =0:14654 The comparison between the actual behavior and the regressed curve by equation (15) is represented in Figure 10 where the real behavior and the regressed one are compared. It has to be highlighted that both that function type (analytical or regression function) and its degree is an absolutely arbitrary choice. Description, implementation, and validation of the simplified model The result of previous analyses is a reduced sdof steering model which can be expressed through equation (16) and graphically synthetized in Figure 11. Mrack \u20acx+ c _x+Kspring x=Finput FESF M \u00f016\u00de Finput =KTBarLinear xinput x \u00f017\u00de with x, _x, \u20acx respectively displacement, velocity and acceleration of the rack, Mrack mass of the rack, c coefficient associated with the damping force, Kspring linear stiffness characteristic of the test bench in opposition to the motion of the rack, Finput force applied to the rack due to the rotation of the steering wheel and FESF M =FESF+FM is the friction force implemented on the basis of what was discussed in paragraph 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure9-1.png",
        "caption": "Fig. 9 Expert based design optimisation (EBDOP-2)",
        "texts": [
            " One alternative to the first proposal design involves reducing only the depth by 10\u00a0mm compared to baseline. This change causes a mere 0.2\u00a0kg weight reduction per part. This proposal design is referred to as SBDOP in this paper. The EBDOP-1 design is shown in Fig.\u00a08. In a second alternative design, some cut-outs are made on the upper and lower panels, but the panels\u2019 thicknesses remain the same. This proposal design is referred to as EBDOP throughout this paper. The EBDOP-2 design is shown in Fig.\u00a09. During the optimisation, 70 iterations have been carried out to obtain the optimised design. The number of iterations used is related to the convergences of the problem, such as changes in the object function or the minimisation of mass. The only gauge change was a reduction of the lower panel thickness from 1.5 to 1.0\u00a0mm. The panel thickness reduction is not a direct outcome of topology or topography optimisation, but it is an outcome of gauge optimisation following the interpretation of a design from the topology and topography optimisation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure10-1.png",
        "caption": "Fig. 10. Method of workspace determination.",
        "texts": [
            " San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at The workspace of the individual R\u0304RR leg i is represented by reachable circle whose radius is lp1 + ld1. The center of each individual reachable circle is initially at the corresponding base platform coordinates A1, A2, A3 (Fig. 9). For finding the mutual workspace of all the three legs, the individual workspace of each leg is translated from their base platform to a vector vi of magnitude equivalent to the radius of the mobile platform Rtop, along the direction of mobile platform orientation \u03d5\u25e6 (Fig. 10). The void circle and the reachable circle (of leg 2) are shown as red dotted circles, before the translation and black circles, after the translation by a magnitude Rtop. The reachable circles C1,C2 and C3 corresponding to each leg are translated, and the obtained mutual workspace is denoted as M1,M2 and M3 in Fig. 11. The mutual workspace area is approximated to be a circle of radius Rw, passing through the intersection points P1, P2 and P3 (Fig. 11). The intersection points are indeed the points of intersection of the medians (of base triangle) with the individual reachable circles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001453_00207721.2020.1803438-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001453_00207721.2020.1803438-Figure1-1.png",
        "caption": "Figure 1. Simple illustration of a small-scale helicopter model.",
        "texts": [
            " , xn) denotes the diagonal matrix with diagonal elements x1, x2, . . . , xn. This section presents the nonlinear dynamic model of a small-scale unmanned helicopter. The helicopter is considered as a six-degree-of-freedom rigid body model. First, two reference frames are defined as follows: (1) earth reference frame (ERF) I = {Oxyz}, which is fixed to the earth; (2) body reference frame (BRF) B = {Obxbybzb}, whose origin is located at the helicopter centre of gravity (Alvarenga et al., 2015; Cai & Chen, 2011). Figure 1 shows the direction of the two reference frames. The dynamic model of the small-scale unmanned helicopter can be described as follows: P\u0307 = V (1) V\u0307 = ge3 + 1 m R( )F (2) R\u0307( ) = R( )S(\u03c9) (3) J\u03c9\u0307 = \u2212\u03c9 \u00d7 J\u03c9 + M (4) where P \u0394= [ x y z ]T \u2208 R 3 and V \u0394= [ u v w ]T \u2208 R 3 refer to the helicopter\u2019s position and velocity vector in the ERF, respectively; m is the helicopter mass, and g is the gravitational acceleration; J represents the approximate inertiamatrix. ThematrixR( ) denotes the rotation matrix from BRF to ERF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000312_1464419319889158-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000312_1464419319889158-Figure3-1.png",
        "caption": "Figure 3. Four types of vibration modes in star herringbone gear transmission system: (a) central member axial\u2013torsional coupling mode; (b) star gear compound mode; (c) central component lateral mode; (d) star gear and sun gear coupling mode.",
        "texts": [
            " The subscripts L and R denote the base-circle helix angles, b, of the matrix for opposite directions. Classification of vibration modes Table 3 lists the natural frequencies and modes of the system, which were calculated from the data in Tables 1 and 2 (the data in Table 2 were derived from experimental tests). Only the first nine natural frequencies are given for the modes with multiplicity m\u00bc 1. The system modes were classified into four categories by analyzing the natural frequency and its vibration modes. A diagram of the vibration modes is shown in Figure 3, where the black dots show the center points of the central members. 1. There are 18 central member axial-torsional coupling modes (multiplicity m\u00bc 1). The frequencies are listed in the first column of Table 3, and the vibration mode is shown schematically in Figure 3(a). The vibration modes of the center member have the following characteristics: the coordinate of the lateral mode of the center element is zero; only axial and torsion modes exist; the vibration modes of the star gears are identical; and the vibration modes of the left and right sides of the herringbone teeth have no regular patterns. 2. The star gears have eight compound modes (multiplicity m\u00bcN-3). These modes appear only for N> 3. The natural frequencies are independent of the number of star gears. The frequencies are listed in Table 3, column 4, and the vibration mode is shown schematically in Figure 3(b). The vibration modes have the following characteristics: the central components do not move; the vibration modes of the star gears are in proportion; and the vibration modes of the left and right sides of the herringbone gears are not regular. Kmesh \u00bc X \u00f0K1 cpN\u00deL \u00fe\u00f0K1 cpN\u00deR 8< : 9= ; 0 0 \u00f0K2 cp1\u00deL \u00f0K2 cpN\u00deL 0 0 0 \u00f0K2 cpN\u00deR P \u00f0K1 rpN\u00deL 0 \u00f0K2 rp1\u00deL \u00f0K2 rpN\u00deL Kr 0 0 0P \u00f0K1 spN\u00deL \u00f0K2 sp1\u00deL \u00f0KN spN\u00deL 0 Ks 0 0 \u00f0K1 pp\u00deL 0 0 0 Kp 0 . . . .. . .. . .. . .. . .. . \u00f0KN pp\u00deL 0 0 0 KpP \u00f0K1 rpN\u00deR 0 \u00f0K2 rp1\u00deR \u00f0K2 rpN\u00deRP \u00f0K1 spN\u00deR \u00f0K2 sp1\u00deR \u00f0K2 rpN\u00deR \u00f0K1 pp\u00deR 0 . . . 0 \u00f0KN pp\u00deR 2 6666666666666666666666666666664 3 7777777777777777777777777777775 \u00f09\u00de KL \u00bc 0 0 0 0 0 0 0 0 0 Kr 0 0 0 0 0 0 0 Ks 0 0 0 Ks 0 0 Kp 0 0 0 Kp 0 . . . .. . .. . .. . .. . .. . Kp 0 0 0 Kp Kr 0 0 0 syms Ks 0 0 Kp 0 . . . 0 Kp 2 666666666666666666666664 3 777777777777777777777775 \u00f010\u00de 3. There are seven central component lateral modes (multiplicity m\u00bc 2). The frequencies are listed in Table 3, column 2, and the vibration modes are shown schematically in Figure 3(c). The vibration modes are characterized as follows: the torsional vibration coordinate of the center component axis is zero, and only lateral motion exists; the vibration coordinates of the left and right sides of the herringbone gears are opposite for the z-axis and the same for the other axes. 4. There are six star gear and sun gear coupling modes (multiplicity m\u00bc 2). The frequencies are listed in Table 3, column 3, and the vibration mode is shown schematically in Figure 3(d). The vibration mode is characterized as follows: all of the central components undergo lateral motion except the planetary cage, which does not move; the vibration modes of the star gears are in proportion; and the vibration modes of the left and right sides of the herringbone gears are the same for the z-axis, and the opposite for the other axes. The first three modes are similar to the modes summarized by Parker for straight gear planetary gears,2 while the fourth category is a new type that arises as a result of the increase in the number of degrees of freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002424_s43154-021-00043-8-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002424_s43154-021-00043-8-Figure7-1.png",
        "caption": "Fig. 7 The gadget for constructing the time-expanded graph. This replaces every undirected graph edge (xu, xv). Traversing a graph edge between steps t and t+1 makes a robot cross the corresponding gadget",
        "texts": [
            " The typical complexities of motion planning here reduce to an optimization problem on a graph. This problem was reduced [16\u2022\u2022] to a multi-commodity network flow by thinking of each robot as a commodity, edges as channels with capacity 1 at any time step. The consideration of discrete time-steps as a dimension for coordinating the robots leads to a time-expanded graph, where copies of vertices are created for every time-step, with a specific edge-structure connecting time step t to time step t + 1. This is shown in the gadget in Fig. 7. The network flow problem is formulated as a mixed integer linear program to optimize the solution in terms of edge weights on the graph. This line of work has extensively applied this method large multi-agent problem instances. Pebble-Graphs Extensive work has been done on an abstraction of this problem [16\u2022\u2022, 29\u201332] which thinks of robots as pebbles on commonly shared graphs. Several algorithmic and complexity results have been demonstrated through the years in this domain. Road Networks A related problem is multi-agent path finding over road networks [33\u201335]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000432_cvci47823.2019.8951669-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000432_cvci47823.2019.8951669-Figure2-1.png",
        "caption": "Figure. 2 Full vehicle model and suspensions in ADAMS [5]",
        "texts": [
            " The four shafts of the power-split gear set are, respectively, connected to the MG1, MG2 engine, and output shaft. In this way, the relationship between the speed of vehicle and the engine can be decoupled. The engine can work more efficiently during the hybrid driving mode through the coordination control of the torques and rotational speeds of the MG1 and MG2. In addition, the electric continuous various speeds (E-CVT) and fuel economy as well as the driving force can be satisfied. III. MODELLING AND SIMULATION Fig.2 demonstrates the dynamic model of the hybrid powertrain, including the engine, engine mounts, driveline system, suspensions and the vehicle body. According to the connect relationship of the power-split device, the speed of carrier and ring gear can be written as 02 1 01 2 02 01 S S PC i n i nn i i (1) 02 1 01 2 01 02 (1 ) n (1 ) nS S R i in i i (2) where PCn is the rotational speed of carrier Rn is the rotational speed of the ring gear; 1Sn is the rotational speed of MG1 2Sn is the rotational speed of MG2 01i is the gear ratio between the ring and small sun gear 02i is the gear ratio between the ring gear and big sun gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure15-1.png",
        "caption": "Fig. 15. Ring displacement in 4PBB.",
        "texts": [
            " 14, fiO and feO are the curvature center of the first contact pair, and siO and seO are the curvature center of the second contact pair, which, respectively, satisfy the geometric relationship defined in Eqs. (6)-(8). Using the same method as in the DGBB, assuming that the displacement of the outer ring in five directions is x\u03b4 , y\u03b4 , z\u03b4 , y\u03b8 , z\u03b8 the ring displacement is decomposed to each steel ball azimuth position, and the raceway center of the outer ring will be move from feO , seO to ' feO , ' seO as shown in Fig. 15. Using the same method as in DGBB, the radial distance ' 1fd , ' 1sd and the axial distance ' 2fd , ' 2sd of the two contact pairs are obtained, as shown in Fig. 16. Note that the two rows of outer raceways in Fig. 16 are located on the same ring, so their radial displacements are equal. But for axial displacement it is mirror-symmetrical, the axial displacement values have the opposite sign. See Eq. (24). ' 1 1 ' 2 2 ' 1 1 ' 2 2 cos sin cos sin f f y i z i zyi yzi f f x zxi yxi s s y i z i zyi yzi s s x zxi yxi d d d d d d d d \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b4 \u03c6 \u03b4 \u03c6 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03c6 \u03b4 \u03c6 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u23a7 = + \u2212 + + \u23aa \u23aa = + + +\u23aa \u23a8 = + \u2212 + +\u23aa \u23aa = \u2212 \u2212 \u2212\u23aa\u23a9 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000249_icems.2019.8921452-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000249_icems.2019.8921452-Figure1-1.png",
        "caption": "Fig. 1. Configurations of DPME motors. (a) Straight-tooth-stator (Existing). (b) Hybrid-stator (Proposed).",
        "texts": [
            " Essentially, DPME motor is a new kind of permanent magnet vernier motor (PMVM), which works based on the bi-directional field modulation effect (BFME). For PMVMs, a straight-tooth stator with open-slots [1, 2], or a split-tooth stator [3, 4] is usually employed to achieve the field modulation effect (FME) [5], wherein the straight teeth or split teeth serve as field modulation teeth in magnetic gears. Structurally, DPME motor has two sets of PMs, one is stator PMs and the other is rotor PMs. As shown in Fig. 1(a), for the existing DPME motors from recent literatures [6-12], all of them adopt straight-tooth stators with open-slots. Both the stator and rotor PMs must be inserted into the stator and rotor open-slots, thus each PM or PM arrays in an open slot and its adjacent iron tooth forms a pair of magnetic poles in order to guarantee the effective coupling between the magnetic field excited by the armature windings and those excited by the stator and rotor PMs. It has been proved that the torque density can be significantly increased by artfully inserting PMs into the stator and rotor open-slots [13, 14]",
            " 2(a) and 2(b) are uniformly distributed and staggered along the circumference, a hybrid stator configuration in Fig. 2(c) can be obtained. And the stator combining a straight-tooth stator and a splittooth stator is called hybrid stator. This work was supported in part by Macao Science and Technology Development Fund (FDCT) of Macao SAR Government under Grant FDCT/040/2017/A1, and in part by the Science and Technology Innovation Committee of Shenzhen under Projects ZDSYS201604291912175. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE Fig. 1(b) shows the configuration of the proposed HSDPME motor. It consists of a hybrid stator and a rotor. The stator PMs are surface-mounted under the straight and split teeth, and each two adjacent PMs are either separated by a straight tooth or by a split tooth. The rotor PMs are inserted into the rotor slots. Both the stator and rotor PMs are magnetized in a unified direction. Thus, the pole-pair number (PPN) of the rotor PMs denoted by p1 is the same as the number of the rotor teeth. The PPN of the stator PMs is denoted by p2, equals three-quarters of the number of stator slots Zs for laying out the armature windings. In order to reduce the consumption of PMs of the existing STS-DPME motors, the proposed motor can be regarded as an evolution of the existing one in Fig. 1(a). Firstly, the rotor remains unchanged, half of the straight teeth are retained and the other half of the straight teeth are replaced with split teeth. Then, stator PMs are surface-mounted under the straight and split teeth, and each two adjacent PMs are either separated by a straight tooth or by a split tooth. Finally, the proposed motor is obtained. Apparently, the proposed motor has reduced the number of stator PMs. In order to mount PMs in the proposed motor, a special design for split tooth, straight tooth and rotor tooth is illustrated in Fig",
            " It can be observed that when p1 is located in the yellow region, the proposed machine can offer a relatively larger back-EMF and output torque, and relatively small cogging torque. So for the proposed machine, the ratio of p2 and p1 should be designed near to 1 for achieving better performances. IV. PEFORMANCE ANALYSIS AND COMPARISON In this section, comparative analysis is carried out by using 2D FEM to verify that the HS-DPME motor can reduce the usage of PMs in existing STS-DPME motor, and make its performances comparable to that of the existing one. As shown in Fig. 1(a), a 3-phase STS-DPME motor with 12- stator-slots and 10-rotor-teeth is selected as a comparative object. According to the evolution in Part B of Section II, the HS-DPME motor in Fig. 1(b) is obtained. Key specifications of the two motors are tabulated in Table III. It is worth noting that the electrical load, copper losses, usage of rotor PMs and key dimensions remain the same. Compared with the existing STS-DPME motor, the volume of stator PMs in the HS-DPME motor is reduced by 51.35%, and the total volume of PMs is reduced by 21.10%. Fig. 7 presents CTs of the two motors. Peaks of their CTs are presented with flags. Fig. 8 exhibits the waveforms of PM flux-linkages in an electric cycle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure7-1.png",
        "caption": "Fig. 7. Simulation model of 2UPR-PRU PKM.",
        "texts": [
            " (31) The actuation forces of all the limbs can thus be obtained as: 3 3 1 1 1 , , , , 1 1 3 3 2 2 2 , , , , 1 1 3 3 3 3 3 , , , , 1 1 ( ) ( ) ( ) = = = = = = \u23a7 \u239b \u239e = \u2212 \u0394 \u2212 \u0394\u23aa \u239c \u239f \u239d \u23a0\u23aa \u23aa \u239b \u239e\u23aa = \u2212 \u0394 \u2212 \u0394\u23a8 \u239c \u239f \u239d \u23a0\u23aa \u23aa \u239b \u239e\u23aa = \u2212 \u0394 \u2212 \u0394\u239c \u239f\u23aa \u239d \u23a0\u23a9 \u2211 \u2211 \u2211 \u2211 \u2211 \u2211 m m m m m m m m m m m m p C p C ij C ij C i j p C p C ij C ij C i j p C p C ij C ij C i j F F F F $ F $ F $ F $ F $ F $ . (32) 4.4 Comparison of theoretical dynamic model with ADAMS model Validation of the analytical dynamic model of 2UPR-PRU PKM is necessary before evaluating the dynamic performance. Here, simulation using ADAMS is carried out to verify the dynamic model, as shown in Fig. 7. The lengths of link parameters are defined as l1 = 135 mm, l2 = 400 mm, l3 = 172.5 mm, and l = 550 mm. The inertia and location parameters of limbs and moving platform of the 2UPR-PRU PKM are listed in Table 1. Considering the actual machining requirements, the polynomial motion trajectory in this simulation which can express the motion of the 2UPR-PRU PKM with three DOFs is set as: ( ) ( ) ( ) 5 5 4 4 3 3 5 5 4 4 3 3 5 5 4 4 3 3 20 6 15 +10 +400 10 6 15 +10 180 10 6 15 +10 180 \u23a7 = \u2212 \u23aa\u23aa = \u03c0 \u2212\u23a8 \u23aa = \u03c0 \u2212\u23aa\u23a9 o d d d d d d d d d z t t t t t t t t t t t t t t t t t t \u03b2 \u03b3 (33) where dt denotes the simulation time, set as 10 s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000703_s12206-020-0122-7-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000703_s12206-020-0122-7-Figure1-1.png",
        "caption": "Fig. 1. Four basic types of threaded fasteners [1]: (a) Screw; (b) bolt and nut; (c) stud and nut; (d) threaded rod and nut.",
        "texts": [
            "1007/s12206-020-0122-7 Keywords: \u00b7 Fatigue test \u00b7 Finite element analysis \u00b7 Clamping force \u00b7 Loosening load \u00b7 Equivalent model Correspondence to: Hoyong Lee hylee@krri.re.kr Citation: Hong, H., Lee, H., Jeong, N., Baek, K., Suh, M. (2020). A study on an equivalent model of the threaded fasteners in complex structures through tightening and loosening analysis. Journal of Mechanical Science and Technology 34 (3) (2020) 1195~1205. http://doi.org/10.1007/s12206-020-0122-7 Received July 24th, 2018 Revised July 3rd, 2019 Accepted December 9th, 2019 \u2020 Recommended by Editor Seungjae Min 1. Introduction Fig. 1 shows four basic types of threaded fasteners. Screws and bolts are by far the most common types, and the difference between them is only one of intended use. Bolts are intended for use with nuts; screws are intended for screwing into tapped holes. Sometimes screws are supplied with the captive washers(usually a lock washer) under the head. A stud is threaded on both ends and is usually screwed permanently into a tapped hole. Threads on the two ends may or may not be identical. A Threaded rod is the least common type"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002074_5.0031388-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002074_5.0031388-Figure3-1.png",
        "caption": "FIG. 3. Schematics of a balloon module for this study. The module consists of a membrane clamped between two circular plates, each of which has a circular orifice at the center, and connected to a ball valve. The balloon is made from inflating a membrane of undeformed thickness t pinned at a circle of radius rb. Once inflated, the balloon contains water of density \u03c1, viscosity \u03bc, volume Vb, and gauge pressure Pb and is subject to gravitational acceleration g. The inflated balloon is approximated as a capillary spherical cap with pinning angle \u03b1b, radius of curvature R, and effective surface tension \u03c3b. The resonance of the balloon can be considered as undamped natural oscillation of a harmonic oscillator with stiffness K, mass m, and displacement y.",
        "texts": [
            " II A. Hydrodynamic and hydroelastic analyses are conducted based on these parameters to reveal the significance of different forces. Phys. Fluids 32, 124113 (2020); doi: 10.1063/5.0031388 32, 124113-3 Published under license by AIP Publishing Subsequently, the experimental techniques for preparing balloons are presented in Sec. II B, and those for oscillating and visualizing balloons are discussed in Sec. II C. Finally, Sec. II D reports the methods for the experiments with drops. As depicted in Fig. 3, a balloon module consists of a membrane clamped between two circular plates, each of which has a circular orifice at the center, and the lower plate is connected to a valve. The membrane is originally flat and uniformly thick. It becomes a balloon when it is subject to hydrostatic pressure from below. The ball valve seals the water in the balloon. The orifice of the upper plate is 120\u25cb countersunk to avoid interfering the balloon\u2019s motion. The balloon is pinned at the circle of minimal radius rb (pinning radius) of the upper plate",
            " In sum, surface force dominates the dynamics of the balloons in this study, and the balloons resonate according to how surface and inertial forces compete. The effects of viscosity and gravity are minor. Accordingly, one should expect the resonance of balloons to resemble that of microliter-sized sessile water drops.54,79 The significance of Bof for the resonance of balloons deserves further discussion. Again, surface and inertial forces dominate the resonance of balloons. Therefore, the motion can be considered as undamped natural oscillation of a harmonic oscillator with mass m and stiffness K, as depicted in Fig. 3. Let y be the displacement and a be the acceleration of the mass. Balancing the forces gives ma + Ky = 0. (6) For the linear oscillation of balloons with the same pinning angle, m is proportional to \u03c1r3 b , a to ab, K to \u03c3b, and y to the deflection yb at the balloon\u2019s apex. Substituting these into Eq. (6) gives \u03c1abr2 b \u03c3b \u221d yb rb , (7) Phys. Fluids 32, 124113 (2020); doi: 10.1063/5.0031388 32, 124113-4 Published under license by AIP Publishing where the left-hand side is Bof and the right-hand side yb/rb \u2261 y\u0304b is a normalized deflection of a resonating balloon"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002010_j.tws.2020.107334-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002010_j.tws.2020.107334-Figure4-1.png",
        "caption": "Fig. 4. Schematic diagram of the screw-type actuator.",
        "texts": [
            " In order to prevent the relative motion between the actuator and the cylindrical shell, several axial sliding blocks and circumferential sliding blocks are used to limit its movement in axial and circumferential Fig. 10. The SMA driving tube. Y. Lu et al. Thin-Walled Structures 159 (2021) 107334 directions. The sliding blocks are fixed on the cylindrical shell by bolts to ensure that the screw-type actuator is closely attached to the inner wall of the shell. The total configuration of the screw-type actuator is illustrated in Fig. 3. Fig. 4 gives the schematic diagram of the screw-type actuator. One can find that the actuator is mounted inside the cylindrical shell. The driving sources of the actuator are SMA tubes with the size of \u00d89mm \u00d7 19.7 mm \u00d7 2.5 mm. The length of the SMA tube after compression in martensite state is 19.04 mm. When heated in unconstrained state, it can recover to 19.64 mm, and the recovery stroke is about 0.6 mm. Fa denotes the driving force generated by the SMA tube. When both ends of the SMA tube are restrained, a single SMA tube can generate a driving force of (0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001378_s11668-020-00961-3-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001378_s11668-020-00961-3-Figure7-1.png",
        "caption": "Fig. 7 Stress distribution of bolt pre-tightening",
        "texts": [
            " Clamping regions of U-bolts are mainly affected by bolt pre-tightening forces. Figure 6 is the clamping structure diagram of the U-bolt in the middle of the leaf spring. The structure is composed of the U-bolt, the spring body, the cover plate, the up plate and the down plate. The spring body is made of composite materials, while other parts are metal. Different U-bolts are selected according to different vehicle loads. Bolts with different grades and diameters have different torques. Standard bolt torques are shown in Table 3. Figure 7 shows the stress distribution of various parts in the leaf spring after U-bolt pre-tightening. Under bolt pretightening deformation, local stress concentration occurred at the four angles where the cover plate and bolts contact. The cover plate and the up plate mainly fix the spring and scatter bolt stress. With the help of the cover plate and the up plate, the stress concentration on the spring body will be significantly reduced. Figure 8 shows the stress distribution at the ends of the composite leaf spring under general working conditions (vertical, braking and steering)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000978_iscid.2019.10108-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000978_iscid.2019.10108-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of rotor rotation with mass eccentricity.",
        "texts": [
            " PRINCIPLE OF ROTOR UNBALANCE VIBRATION IN The suspension force and torque of a BSRM for A-phase can be derived as [8]: 1 1 2 2 2 1 1 2 x ma sa sa y ma sa sa F i K i K i F i K i K i (1) 2 2 2 2 2 2 1 1 2 22t m ma s sa s saT K N i N i N i (2) where ima and Nm are the current and turns of main winding; Fx and isa1 are the force and current of x-axis suspension winding; Fy and isa2 are the force and current of y-axis suspension winding; Ns is the suspension winding turns; K1 and K2 is the x- and y-axes suspension force coefficient, Authorized licensed use limited to: University of Canberra. Downloaded on June 07,2020 at 16:10:42 UTC from IEEE Xplore. Restrictions apply. respectively [8]; Kt is the electromagnetic torque coefficient. The expression of suspension force of B- and C-phase has resemblance to that of A-phase. Fig. 1 shows the schematic diagram of rotor rotation with mass eccentricity. When the geometric center O does not coincide with the mass center C, the unbalance magnetic pull will be produced. Because the centrifugal force caused by mass eccentricity is proportional to the squared rotor speed wn, the centrifugal force will increases with the increment of the speed, and the unbalance vibration will become more severe. III. UNBALANCE VIBRATION COMPENSATION CONTROL IN BSRM BASED ON EKF The rotor dynamic equations for a BSRM are expressed as: 2 sin 60 2 cos 60 2 60 r u r u x r y r n u n u t n L x x x y y y F mx F my mg w x A t w y A t j T w T (3) where xu, xr, x and yu, yr, y are the unbalance vibration displacement, the random displacement and the actual displacement in x- and y-axes, respectively; A is the magnitude of unbalance vibration signal; TL is the load torque; jt is the rotational inertia; m and g are the rotor mass and the gravity acceleration, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure2-1.png",
        "caption": "Fig. 2 Four metamorphic configurations of the 2-DOF metamorphic mechanism",
        "texts": [
            "\u00a01b, d) and its composition principle are shown in Fig.\u00a01. The mechanism has two loops. The active part and frame are connected with augmented Assur group RRRR to form a loop. The other loop is composed of Assur group RRP connected to the former loop and frame. The metamorphic process of mechanism based on augmented Assur group is the process of transforming 1-DOF augmented Assur group into basic Assur group. The 2-DOF metamorphic mechanism shown in Fig.\u00a01 consists of an augmented Assur group, and its four metamorphic configurations are shown in Fig.\u00a02. According to the above-mentioned structural theory and formation methodology of metamorphic mechanisms based on augmented Assur groups, the unified dynamics modeling method of constrained metamorphic mechanism is studied. The driving forms of active parts are shown in Fig.\u00a03. Figure\u00a03a and b shows active parts in pure rotational and pure prismatic form, respectively. According to the Newton\u2013Euler equation (referred to as N/E equation), the dynamic equations of active parts can be written as follows, where Fi\u22121,i and Fi+1,i are the force vectors at component Li exerted by components Li\u22121 and Li+1 , respectively, and (1) { Fi\u22121,i \u2212 Fi,i+1 + Fi = miC\u0308i Mi\u22121,i \u2212 li \u00d7 Fi\u22121,i \u2212Mi,i+1 \u2212 hi \u00d7 Fi,i+1 +Mi = IC,i\ud835\udf3ai Fi+1,i = \u2212Fi,i+1 ; Mi\u22121,i and Mi+1,i are the moments at component Li exerted by components Li\u22121 and Li+1 , respectively, and Mi+1,i = \u2212Mi,i+1 ; Fi is the external force vector acting on the component Li ; Mi is the external moment acting on the component Li ; IC,i is the moment of inertia of the component Li around the centroid Ci ; i is the angular acceleration of the component Li ; C\u0308i is the acceleration vector at the centroid of the component Li"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001125_j.matpr.2020.05.084-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001125_j.matpr.2020.05.084-Figure3-1.png",
        "caption": "Fig. 3. Effects of changing the values of L/2R ratios on the Von Mice\u2019s stress distribution",
        "texts": [
            " The structural deformations were simulated by finite element analysis for linear material conditions, choosing stainless steel 316L as the base material. The uniform FEA mesh of the tetrahedron shape and 0.5 mm of size was generated for all the models. Fig. 2 is depicting the applied boundary conditions, and image in the inset shows the tetrahedral discretisation more closely for horizontal elliptical form. The size of the finite element mesh at around 0.5 mm was obtained based on a mesh convergence analysis result indicating the lateral displacement to converge on a fixed value when the element size was iteratively reduced to around 0.50 mm. Fig. 3 shows the stress distribution patterns and auxetic responses emanating in the three different S-shaped model variants analysed. The stress distribution patterns in Fig. 3 (a), (c) and (e) are the results obtained by changing the geometrical parameter L/2R. With the change in the structural form from circular shape (Fig. 3(a)) to the horizontal elliptical form (Fig. 3(c)), a noticeable reduction in the stress intensity from 390 MPa (Circular shape) to 260 MPa (Horizontal elliptical) is obtained. However, when the structural form was changed to the vertical elliptical form in Fig. 3(e), the stress intensity again increased and reached almost up to 380 MPa, equivalent to the stress intensity of the circular model (Fig. 3(a)). The variations in the stress intensities of these two structural forms (Horizontal and vertical elliptical forms) of the circular S-shape clearly elucidated that the varying structural forms adopt different deformation modes as a result of the changes in the values of the L/2R values. The dark red areas in Fig.\u2019s 3 (a), (c) and (e) are indicating the stress concentration zones. Higher stress zones in the case of the circular (Fig. 3(a)) and the vertical elliptical (Fig. 3(e)) S-shapes are concentrated towards the centre. However, the horizontal elliptical form of Fig. 3(b) shows the stress concentration away from the central portions. Evidently, the stress concentration zones as well as the stress intensities varied with the geometrical varia- ipse and (c) Vertical ellipse S- shaped unit cells. Please cite this article as: K. Meena and S. Singamneni, An elongated S-shaped a https://doi.org/10.1016/j.matpr.2020.05.084 tions in the S-shaped unit cells. More significantly both stress concentration zones and the intensities attained more favourable modes in the case of horizontal elliptical S-shape form due to changes in the L/2R ratio. The lateral deformation results of all the three models, with changing L/2R ratios are shown in Fig. 3(b), (d) and (e). It may clearly be noted that the deflection and deformation patterns had negligible variations due to the changing geometrical shapes. The maximum deformation levels, the deformation distribution patterns as well as the deflections in the overall shapes of the unit cells remained the same from the circular S-shape form to the horizontal and vertical elliptical S-shape forms. This is indicative of the fact that the favourable stress distribution was achieved at no loss of the deflection responses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001428_tec.2020.3017103-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001428_tec.2020.3017103-Figure5-1.png",
        "caption": "Fig. 5. PM pole with sawtooth edge and integral diagrams of force and moment. (a) Diagram of PM pole. (b) Diagram of radial force and bending moment integral path.",
        "texts": [
            " But, the second harmonic component of the slot frequency vibration is not significantly reduced. Optimization of the shape of PM poles should be go further to deal with all harmonic components at the same time, which is easy to execute in practice. This is the object of later sections. The analysis process of SEP scheme is similar to that of PVPW scheme. Extending the idea and method of PVPW [18], the step number should be increased to mitigate all the harmonic components of slot frequency vibration. So the PM pole with the sawtooth edge is proposed and shown in Fig. 5 (a). In axial direction, the PM pole is divided into 8 segments, where the width of every segment is continuously changed and the edge is sawtooth shape. Fig. 6 shows the photo of the PM pole with the sawtooth edge, where the whole pole is divided into two pieces in axial direction to facilitate manufacturing, and there are two invalid segments without sawtooth in both two ends. The two invalid segments are positioned outside the rotor core. The rotor of this motor is exactly the same as that of Fig. 1 (b). Fig. 5(b) shows the size of one of the sections,  refers to the number of tooth pitch that is corresponding to the rotor at the narrowest circumferential direction of the magnetic pole, and   denotes the number of tooth pitch that corresponding to the rotor at the magnetic pole with the widest Authorized licensed use limited to: University of New South Wales. Downloaded on September 20,2020 at 02:17:15 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. length. The radial force acting on the PM pole can be obtained by integrating the force density on the surface. In accordance with symmetry, integral of segment 1 is conducted according to Fig. 5(b) firstly. The integral path is to integrate firstly in the circumferential direction and then in the axial direction. The total radial force is obtained by adding the integral results of 8 sections 1 2 0 1 0 2 ( ) 8 1 8 [ cos( )sin( )sin ] 2 2 l C np nkC np k r k p t p Rd dx p k k kZj kZ t k k p               (5) where, 2 00 1 04 /npp B Rl Z    , 1 12 2C j p Z x Zl     , 2 12 2C j p Z x Zl     . Under the effect of rotor rotation, the radial force on both sides of the axial centerline of magnetic pole changes with time, which forms the bending moment on the centerline of magnetic pole. The calculation method employed for bending moment is shown in Fig. 5(b). Moreover, the method of double integral is adopted for the calculation of the bending moment. It can be carried out by firstly integrating along the circumferential direction, and then integrating along the axial direction, followed by finally summing up the integral results of the 8 sections, so as to obtain the total bending moment as 1 2 0 1 ( ) 8 sin l C np n cC m t p R Rd dx            2 2 08 A B 1 C D 1knp k m kZ kZ              (6) where, 0 2 0 1 02np cm B R R Zl   , Rc is the radius of the outer surface of the pole"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002629_012013-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002629_012013-Figure5-1.png",
        "caption": "Figure 5. Motor housing to end cover flange",
        "texts": [],
        "surrounding_texts": [
            "Modal Analysis is mostly dispensed to seek out natural frequency of the system and completely different mode shapes at different Eigen values [12,14]. In this work, totally different modes and Eigen values square measure calculates and premeditated between mode shapes versus natural frequencies [15,20, Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 21]. This shows the stiffness and rigidity of the material toward totally different modes. Using Modal analysis it can be captured the first 10 natural frequencies and related mode shape of each frequency for the 3 cases. Figure 7 shows the modal analysis results for case A. Each figure represents the frequency of each mode and displacement related to the frequency. Due to rotational velocity the model is exhibiting rotational vibration. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Table 1 shows the value of modal analysis for case A, B and C. According to this table Case C is having more natural frequency than other 2 cases. Figure 8 shows the relation between modes and damped frequency. Based on modal analysis results Case C (Hybrid Hemp - Flax fibre) has additional stiffness (due to high natural frequency) while compared to other two cases. Conclusion In this analysis, a study for three totally different composite materials particularly Hemp fibre, Flax fibre and Hybrid of Hemp and Flax fibre for Finite Element Analysis of electrical motor casing has been carried out. The motor casing should be rigid and better in stiffness to resist operative vibrations. 3D modelling of motor casing is built using Parametric Creo one among the fine 3D modelling software. These 3 materials are applied to motor casing as a material properties of the model and three types of analysis namely, Modal Analysis, Rotor Dynamics and Structural Transient Analysis carried out to check the stiffness, reliability towards operative frequency and durability of motor casing. As determined from modal analysis Case C (Hybrid Hemp Flax fibre) has better stiffness compared to two different fibres as shown in graph fig 7 (Modes vs Damped Frequency). Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Campbell diagram was premeditated between frequencies vs speed of the motor for predicting critical speed. The operative frequency and operative speed do not match with the critical speed and frequency. Model analysis was performed for finding the vibrational stability of the composite materials and observation Case C shows low deflection and equivalent stresses compared to different fibres. Force vs deflection curve and stress vs strain curve are premeditated to indicate the stiffness of the materials and energy vs deflection curve premeditate show of the strain energy observed, when deformation happens. As determined from all above studies Case C has higher stiffness and strength compare from different two fibres. Reference [1] Bajuri F, Norkhairunnisa Mazlan, Mohamad Ridzwan Ishak and Junichiro Imatomi, 2016, \u201cFlexural and compressive properties of hybrid kenaf/silica nanoparticles in epoxy composite\u201d, Procedia Chemistry, Vol.19, pp:955-960. [2] Korniejenko. K, Frczek. E, Pytlak. E and Adamski. M, 2016, \u201cMechanical properties of geo polymer composites reinforced with natural fibres\u201d, Procedia Engineering, Vol.151, pp:388393. [3] Lut Pil, Farida Bensadoun, Julie Pariset and Ignaas Verpoest, 2016, \u201cWhy are designers fascinated by flax and hemp fibre composites?\u201d, Composites: Part A, Vol.83, pp:193-205. [4] Shuhimi F, Mohd Fadzli Bin Abdollah, Kalam M.A and Hilmi Amiruddin, 2016, \u201cTribological characteristics comparison for oil palm fibre/epoxy and kenaf fibre/epoxy composites under dry sliding conditions\u201d, Tribology International 101. [5] Raman Bharath, Vijaya Ramnath B and Manoaran N, 2015, \u201cKenaf fibre reinforced composites: A review\u201d, ARPN Journal of Engineering and Applied Sciences, Vol.10, No.13, pp:5483-5485. [6] Zamria M, Hazizan Md Akil and Zainal Ariffin Mohd Ishak, , 2016 \u201cPultruded kenaf fibre reinforced composites: Effect of different kenaf fibre yarn tex\u201d, Procedia Chemistry, Vol.19, pp:577-585. [7] Bensadoun. F, Depuydt. D, Baets. J, Verpoest. I, Van Vuure. A.W, 2017, \u201cLow velocity impact properties of flax composites\u201d, Composite Structures, Vol.176, pp:933-944. [8] Changlei Xia, Jason Yu, Sheldon Q. Shi, Ying Qiu, Liping Cai, Felix Wu. H, Han Ren, Xu Nie and Hualiang Zhang, 2017, \u201cNatural fibre and aluminum sheet hybrid composites for high electromagnetic interference shielding performance\u201d, Composites Part B, Vol.114, pp:121- 127. [9] Ming Liu, Andreas Baum, Jurgen Odermatt, Jens Berger, Liyun Yu, Birgitte Zeuner, Anders Thygesen, Jesper Holck and Anne S. Meyer, 2017, \u201cOxidation of lignin in hemp fibres by laccase: Effects on mechanical properties of hemp fibres and unidirectional fibre/epoxy composites\u201d, Composites: Part A, Vol.95, pp:377-387. [10] Bambach. M.R, 2018, \u201cGeometric optimisation and compression design of natural fibre composite structural channel sections\u201d, Composite Structures, Vol.185, pp:549-560. [11] Ravi Prakash Babu K., Rao P.K.V., Ali M.A., Raghu Kumar B., 2019, \u2018Crack simulations in shaft-hub-blade system using modal analysis\u2019, International Journal of Recent Technology and Engineering, 8(1), PP.2646-2650. [12] Bala Subramanyam P.N.V., Rao B.N., Prakash R.L., Rajavardhan D.S.K., Sankarudu B.S.K., Ganesh P.,2019, \u2018Modal analysis on go-kart chassis\u2019, International Journal of Engineering and Advanced Technology, 8(4), PP.1701-1705. [13] Syed K., Mannam H., Koya M.K., Meduri S.K., Laxmi V.S., Kasturi S.P., 2019, \u2018Dynamic simulation of gas turbine blade using finite element analysis\u2019, International Journal of Mechanical and Production Engineering Research and Development, 9(), PP.745-751. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 [14] Dama K.K., Surya S., Babu C., Pavan, Kiran R., 2018, \u2018Modal analysis of an automotive subsystem using cae techniques\u2019, International Journal of Mechanical Engineering and Technology, 9(3), PP. 1157-1162. [15] Deepak Kumar G., Kommuri N., Kota H., Prasanth V., Hameed S, 2018, \u2018Vibrational analysis of a human body on an automobile\u2019, Journal of Advanced Research in Dynamical and Control Systems ,10(6), PP. 39-45. [16] Devi S.K., 2018, \u2018The optimization of layer densities for multilayered insulation systems by using ansys\u2019, International Journal of Mechanical and Production Engineering Research and Development ,8(4), PP. 1087-1098. [17] Balasubramanyam P., Nageswara Rao B., Banerjee M., Lakshmana Swamy B., 2019, \u2018Impact analysis on Go-Kart chassis with variable speeds using ansys 19.0\u2019, International Journal of Engineering and Advanced Technology, 8(6), PP.2614-2620. [18] Kancheti N., Reddy Vemula A., Reddy Gudibandla G., Krishna H., Bala Subramanyam P.N.V. 2019, \u2018Modeling and analysis of wheel rim using ansys\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(8), PP.415-418. [19] Bharath E.S., NagaSree Harsha O., Bala Subramanyam P.N.V., Reddy G.S.S., 2019, \u2018Impact analysis on honey comb structured go-kart bumper using ansys R19.1\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(6), PP.396-400. [20] Ravi Prakash Babu K., Gupta S.K., Phaneendra Sai Sri D., Gautam K., Kumar C.M. 2017,\u2019Significance of modal analysis to detect and diagnose misalignment fault in turbine rotor\u2019,Journal of Advanced Research in Dynamical and Control Systems,9(Special issue 14),PP.2637-2648. [21] Babu K.R.P., Prasanth K.F.V., Sai V.V., Kumar M.V.S., Kumar B.N.S., Chaitanya B.S. 2017,\u2019Effect of pre-twist on free vibration characteristics of a cantilever beam with experimental validation\u2019,International Journal of Mechanical Engineering and Technology,8(5),PP:389-399. [22] T. Srinivasan, CH.V.L.N. Ram Gopal, P. Sasidhar, P. Vineeth Kumar and Shaik Sohail, 2020, Effect of change of cross section and Frequency Analysis of Anti-Roll Bar Using FEA, IOP Conf. Series: Materials Science and Engineering 954 (2020) 012027 IOP Publishing, doi:10.1088/1757-899X/954/1/012027"
        ]
    },
    {
        "image_filename": "designv11_63_0001255_2020.08.48-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001255_2020.08.48-Figure4-1.png",
        "caption": "Figure 4. (A) Utilized microfabricated arrays (left) and their associated fluidic housing (right). (B) Schematic representation of the cross-linked functional PEG polymer electrode modification and subsequent antibody integration. Reprinted with permission from reference [52]. Copyright 2014 American Chemical Society.",
        "texts": [
            " Therefore, it is necessary to design an antifouling sensing platform for effectively reducing undesired binding on the electrode surface to maintain biosensor performance in practical analysis. Among plenty of antifouling materials integrated on the electrode surface, poly(ethylene glycol) (PEG) polymer are extensively used due to well biocompatibility, naturally inert and hydrophilicity. Davis\u2019 group integrated cross-linked PEG polymer films generated from commercial PEGylated monomers within fabricated microelectrode arrays for simultaneous detection of insulin and CRP in human serum (Figure 4) [52]. An optimized molar ratio (2:3) of 4-armed PEG-epoxide and PEG-amine was selected for the formation of the thermo-polymerized film, ensuring enough accessible amine groups for antibody attachment and the performance of the resulting biosensor. In the absence of amplifying redox probes, CRP was monitored by non-Faradaic EIS with a linear range of 0.5 ~ 50 nM (R2 = 0.997) and a LOD of 150 \u00b1 10 pM. Cellulose nanofibril not only is biocompatible and biodegradable, but also has unique nanostructures in films with high mechanical strength, small porosity and high density"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001226_acpee48638.2020.9136520-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001226_acpee48638.2020.9136520-Figure9-1.png",
        "caption": "Fig. 9. The diagram of analysis for defects position influnence on the electric field distribution",
        "texts": [
            " The electric field with the defect at various position parameters are shown in Table V, and the length of the defect is 0.5m. When the defect position is close to the high voltage end of the insulator, the electric fields at the end of the defect increase. When the parameter S0 is 0.2m and the defect length is 0.5m, the electric field at the defect can reach to 905.4kV/m. b) Defects in the Both Insulator Strings: The influence of the defect positions on the electric field distribution of the composite insulators is shown in Fig. 9. In Fig. 9, the parameter S01 and S02 are the distances between the defect and the high voltage end of the left and right insulator string. The parameters L1 and L2 are the length of the defects. The defects are at the floating potential. Under this condition, the length and position of defects will affect the calculation results. In order to simplify the calculation, it is assumed that the defect length is the same and the defect is located at the same height of the V-type insulator strings. When L1 and L2 are set to 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000534_iros40897.2019.8968016-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000534_iros40897.2019.8968016-Figure7-1.png",
        "caption": "Fig. 7. MDU 1100T X-Band Doppler motion detection unit experimental setup.",
        "texts": [
            " Since we obtain location information by opening the switch of Fig. 2, a limited controllability and therefore the lack of a VCO circuit can be tolerated in our evaluation prototype. On the receiver side, a DRO is employed also, allowing for a good control of the reference frequency, which provides a stable reference signal at approximately the same operating frequency as the transmitter DRO. As a multiplier, we used the MDU1100T receiver\u2019s single balanced mixer consisting of two Schottky diodes in an antiparallel configuration. As depicted in Fig. 7, the transmitter is moved away from the receiver, using a platform on wheels construction. A foil sheet is prohibiting receiver antenna to transmit and transmitter antenna to receive, in order not to have undesirable signal mixing, into each unit\u2019s mixing stage or interference in general. A measurement of the intermediate frequency (IF) output of the receiver unit is consequently performed, moving the transmitter while maintaining the receiver unit in place. Because both DROs are operating at about the same frequency, the IF signal is in essence centered at dc; therefore, it is a zero intermediate frequency (ZIF) and a common place oscilloscope may be used to monitor it. The first experimental setup measured the transmitter\u2019s displacement in respect to the fixed receiver covering a distance of 8 cm, while the movement is concluded in a 2 s time interval. The transmitter velocity is constant and controlled with the use of a microcontroller controlled step motor. This setup is shown in Fig. 7. Ten measurements were obtained using a digital oscilloscope to capture the IF output in order to extract statistical results and estimate system precision. A typical captured IF waveform is shown in Fig. 8. It is clear that the resultant IF signal is of sinusoidal nature, which is in agreement with theoretical predictions. The phaseshift method analysis of Section III was therefore deemed valid. All waveforms were curve fitted in order to accommodate subsequent noise or interference filtering. The time-domain regression analysis method which was selected is based on locally weighted scatterplot smoothing (LOWESS) [36]",
            " This corresponds to a precision of 3 \u00d7 0.653 mm = \u00b11.959 mm. Since a medical instrument\u2019s displacement is not always performed with constant velocity as evaluated previously, additional setups with different experimental conditions were implemented, in order to evaluate the system\u2019s characteristics in more detail. In this experiment, a setup was formed around a platform on wheels, capable of moving the transmitter along a linear trajectory toward or away from a stationary receiver. This setup was similar to the one shown in Fig. 7 and is designated in Fig. 9. The setup described in the previous subsection measured the system\u2019s transmitter displacement with constant velocity away from the stationary receiver. In order to prove the capability of the system to operate with nonconstant velocities, as the case of an IVUS catheter movement, we did perform an experiment where the transmitter was displaced from the stationary receiver with uniform acceleration. Thus, the setup aims to evaluate the system characteristics in order to verify its capabilities on measuring the exact position of any medical instrument moving inside a human body with whatever speed, either constant or variable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000555_j.jmapro.2020.01.054-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000555_j.jmapro.2020.01.054-Figure8-1.png",
        "caption": "Fig. 8. Three-dimensional finite element mesh for the components fabricated by AM based on substrates of different geometries: (a) BO-3, (b) PL, and (c) GR-3 schemes (the red line represents the interface between forged substrate and additive area). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).",
        "texts": [
            " The Gaussian fitting method was used to locate the peak and corrections were made for the background, absorption factor, and Lorentz-polarization factor. The elastic modulus E and the Poisson\u2019s ratio \u03bd were 110 GPa and 0.31, respectively. According to the shape and size of the actual components (see Fig. 4) in the above three schemes, three coupled thermomechanical models were established in the ANSYS finite element software to simulate the deposition process of the single-pass multilayer walls. The finite element mesh of the components is shown in Fig. 8. The numbers of layers of the models in GR-3, PL, and BO-3 were 180, 150, and 120, respectively. Except for the area near the groove in GR-3 that used a free mesh due to the complex structure, other areas of the three models used a mapped mesh. To improve the accuracy of calculation, a finer mesh was used at the location close to the cladding because the thermal and stress gradients at that location are usually the largest. In the additive area, 8-node hexahedral elements with dimensions of 2\u00d7 2 \u00d7 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001904_icem49940.2020.9270724-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001904_icem49940.2020.9270724-Figure7-1.png",
        "caption": "Fig. 7. CAD model of the machine under investigation for topology Fig. 6e and f. The corresponding coils are denoted by matching colors. The stator coils are subdivided into the six coils of Fig. 3 for each phase and each pole. The double layer rotor winding is according to Fig. 4. The rotor circuit is modelled using a coupled electrical network. The flux lines on the left denote no-load operation of pole pair pmain. On the right, the no-load flux lines if only pharm is excited are shown.",
        "texts": [
            " This reduces the switching losses and conduction losses of the power electronics and simplifies the control. The differential MOSFET pair for the step-down converter is required to ensure a symmetrical load. In order to demonstrate the potential and limitations of harmonic excitation, the topologies according to Figs. 6a, 6e and 6f are used and the efficiencies of the excitation concepts are calculated based on Finite-Element-Analysis (FEA) for a 10 kW non-salient-pole machine according to Tab. I. The geometry of the machine is depicted in Fig. 7. The specifications of the power electronic components used for the rotor circuit are given in Tab. II. Each topology offers a number of degrees of freedom which affect the efficiency. In order to determine the operating points, the following procedure which neglects saturation due to currents of pole pair pharm is used. 1) Calculate the differential coupling inductances between stator and rotor dLfd,harm, the self inductance dLf,harm and the load inductance of the pmain field coil Lf,main. Iron saturation is defined only by the fundamental currents {Iq,main, If,main}",
            " Furthermore, if the machine is saturated, the magnetic d-axis of pole pair pharm (and pmain) shifts from the saturated iron (the geometric d-axis), denoted by coupling inductances Ldq,harm 6= 0, Lqf,harm 6= 0 which can create parasitic pharm torques. It is expected that this also negatively affects the torque. Nevertheless, it can be concluded that the analytical operating point calculation can be used to predict the machine behaviour with satisfactory agreement. It should be noted for the given machine in Fig. 7 that asymmetric pulling forces act on the rotor, although they are not excessive. This is due to the interaction of the pmain = 2 fundamental field and the pharm = 1 subharmonic field. However, this is not an inherent problem with the excitation concept as these asymmetric forces disappear for pole-pair combinations |pmain \u2212 pharm| > 1, e.g. for pmain = 4 and pharm = 2. In this work, an extension to an existing harmonic brushless excitation scheme for WRSMs has been proposed, which enables high starting torque and higher efficiencies than conventional harmonic excitation topologies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure13-1.png",
        "caption": "Fig. 13. On-load magnetic field distribution of CPM-I. (a) 10A, (b) 12A, and (c) 15A.",
        "texts": [],
        "surrounding_texts": [
            "The back-EMF and electromagnetic torque are shown in Fig. 15. It can be observed that although the PM volumes of CPM-I and CPM-II are similar, the electromagnetic performance of CPM-II is much better than that of CPM-I. In addition, the electromagnetic performance firstly increases and then decreases as \u03b4 increases. When \u03b4=60 deg, it reaches the optimal value. This is consistent with the variation of the air-gap flux density as shown in Fig. 10. It can be found that when \u03b4\u226575 deg, the saturation of the salient iron pole can be eliminated. However, when \u03b4\uff1e90deg, the electromagnetic performance is reduced significantly. Hence, considering the saturation of salient iron pole and the electromagnetic performance, an excellent balance between reduced saturation and good electromagnetic performance can be achieved when 75deg\u2264\u03b4\u226490 deg is satisfied. Considering the manufacturability, the machine with \u03b4=90 deg is chosen for the prototype machine."
        ]
    },
    {
        "image_filename": "designv11_63_0002872_j.procir.2021.05.033-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002872_j.procir.2021.05.033-Figure3-1.png",
        "caption": "Fig. 3. Theoretical supplemental surface concept [2]",
        "texts": [
            " There are other unit cells besides struts, but their use involves further singularities that will not be developed here. In fact, strut-based unit cells are the most widely used and are the focus of our present research. Due to the nature of lattice structures, not all surfaces are continuous, which makes the geometric tolerancing and verification difficult. Indeed it can be challenging to determine which element belongs to a functional surface. Thus ASME Y14.46 [2] has provided a response to this problem by introducing the concept of theoretical supplemental surfaces (TSS) (figure 3). They are defined as theoretical surfaces that can be toleranced and associated with a set of real points following the same idea as the tolerancing of integral or derived features [8]. Ameta et al. [1] applied this principle to a real part by defining a derivated supplemental surface, which is a surface obtained from points measured with a confocal laser microscope. The points are recorded on external nodes of planar shape. The surface is then calculated according to two association criteria, least-squares and Chebyshev to check a flatness tolerance",
            " Lattice structure generation on Rhino-Grasshopper Many plug-ins are available for Grasshopper, enabling additional functionalities. Crystallon [4] can be used to generate conformal lattice structures (figure 6). Thus allowing the generation of much more complex shapes than with other solutions. However, the plugin only generates wireframes, so in order to make the volumic structure, it has to be coupled with Dendro [5]. Finally, the model obtained can be exported as STL using the Pancake add-on [19]. K\u00e9vin Ferreira et al. / Procedia CIRP 100 (2021) 846\u2013851 849 K. Ferreira et al. / Procedia CIRP 00 (2019) 000\u2013000 3 Fig. 3. Theoretical supplemental surface concept [2] acteristics restricts the use of conventional measurement instruments. Thus X-ray Computed Tomography (XCT) is often regarded as the most suitable method for observing these structures. The advantage of this technique lies in the fact that it allows retrieving information inside and outside the part and even to have information on the geometric quality of the surface. Furthermore, XCT offers the possibility of aligning the measured data with the 3D CAD model in order to observe the defects globally"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000901_s11012-020-01160-y-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000901_s11012-020-01160-y-Figure4-1.png",
        "caption": "Fig. 4 Geometrical parameters for the fillet-foundation deflection [16]",
        "texts": [
            " (5) can be expressed as follows Ub \u00bc Z q cos a 0 Fb\u00f0dC xAB\u00f0u\u00de\u00de FahC\u00bd 2 4 3 EBy3AB\u00f0u\u00de dxAB\u00f0u\u00de du du \u00fe Z uC q cos a Fb\u00f0dC xBD\u00f0u\u00de\u00de FahC\u00bd 2 4 3 EBy3BD\u00f0u\u00de dxBD\u00f0u\u00de du du Us \u00bc Z q cos a 0 1:2F2 b 4GByAB\u00f0u\u00de dxAB\u00f0u\u00de du du \u00fe Z uC q cos a 1:2F2 b 4GByBD\u00f0u\u00de dxBD\u00f0u\u00de du du Ua \u00bc Z q cos a 0 F2 a 4EByAB\u00f0u\u00de dxAB\u00f0u\u00de du du \u00fe Z uC q cos a F2 a 4EByBD\u00f0u\u00de dxBD\u00f0u\u00de du du \u00f06\u00de where Fa \u00bc F sin b, Fb \u00bc F cos b, aC \u00bc arccos rb rC , c \u00bc p\u00fe4e tan a 2Z \u00fe inva invaC, b \u00bc aC c, uC \u00bc \u00f0h c\u00der sin2 a\u00fe \u00f0q cos2 a\u00fe h sin a\u00de cos a (see Fig. 2), E and G represent elastic and shear moduli of gear tooth, respectively. In addition to the tooth deflection, the filletfoundation deflection, df, also conduces the deformation of the meshing tooth pair. The analytical formula for dfwas derived by Sainsot and Velex [16] base upon the theory of Muskhelishvili [17]. It is given by df \u00bc F cos2 b EB L uf Sf 2 \u00feM uf Sf \u00fe P 1\u00fe Q tan2 b ( ) \u00f07\u00de where parameters Sf and uf are indicated in Fig. 4. The coefficients L*, M*, P* and Q* can be calculated according to the following polynomial functions [16], X \u00f0hf ; hf \u00de \u00bc Ai h2f \u00fe Bih 2 f \u00fe Cihf hf \u00fe Di hf \u00fe Eihf \u00fe Fi \u00f08\u00de where X* indicates the coefficients L*,M*,P* andQ*; rf, rint and hf are shown in Fig. 4; hf \u00bc rf =rint; the values of Ai, Bi, Ci, Di, Ei and Fi are summarized in Table 1. 2.3 Elliptical Hertzian contact for crown gear pairs For gear pairs with circular arc crowning, line contact of tooth surfaces is substituted by instantaneous point contact [18]. Because the contact area is relatively small, two meshing gear teeth can be considered as a half-space. The contact of the tooth pair during meshing can be approximately described as an Elliptical Hertzian contact. Consequently, the contact zone and the normal approach of two elastic bodies can be obtained according to Hertzian law"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001188_tmag.2020.3004840-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001188_tmag.2020.3004840-Figure3-1.png",
        "caption": "Fig. 3. Decomposition of the rotational flux (RD: rolling direction, TD: transverse direction).",
        "texts": [
            " The nonlinearities of the GO and NO steel plates are considered applying the initial B-H curves shown in Fig. 2. The B-H curves of GO steel plates are measured under flux density in arbitrary directions [7]. B. Iron Loss Calculation Method [8] Rotational flux are generated in the T-joints in the threephase reactor, thus the rotational iron loss due to the rotational flux are calculated. First, the rotational flux is decomposed into the maximum flux densities BL and BS along the major and minor axes of the rotational flux as shown in Fig. 3. Then, for NO steel plates, the iron loss Wi,rot is the sum of the alternating iron losses Wi,alt (BL) and Wi,alt (BS) using the iron loss curve in Fig. 4 (a) as follows: , , ,( ) ( )i rot i alt L i alt SW W B W B  (1) The alternating iron losses for a GO material differ with \u03b8L as shown in Fig. 4 (b), in which \u03b8L is the angle of the alternating flux density to the rolling direction as shown in Fig. 3. Therefore, the rotational iron loss Wi,rot for the GO material is the sum of Wi,alt (BL, L ) and Wi,alt (BS, S ) as follows: , , ,( , ) ( , )i rot i alt L L i alt S SW W B W B   (2) The local electromagnetic force fem (ip) at each node ip is calculated as follows [9]: ( ) ( )grad ip ip em N dV  f T (3) where T is the Maxwell stress tensor and N is the shape function. For the NO steel plates, T is expressed as follows: ij i j ijT H B d   B H (4) where the subscript i, j denote the x, y, and z components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002206_00219266.2020.1858927-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002206_00219266.2020.1858927-Figure1-1.png",
        "caption": "Figure 1. (a) Large and small receptors, resembling antibody binding sites and viral epitopes respectively. (b) Adapters of varying size: the small and mini adaptors form the base of IgM and IgA, respectively, while the large adapter forms the base of viruses. (c) Antibody models, made by adding large receptors onto mini and small adapters. (d) Bacteriophage T4.",
        "texts": [
            " The kit contained models of antibodies IgG, IgA, and IgM, three of the most prevalent found in the human body. Although the students were not previously been exposed to specific examples of antibodies in previous lectures, they were given an explanation of their names and roles as they were freely interacting with the prints. This was kept at a basic level and focused solely on the antibody name, role, and basic structure. These antibodies were made from printed models of two adapter-like figures and one receptor-like figure that is forked (Figure 1). The adapter-like figures represented the base of the antibody, while the large receptor connected to the adapter, acting as the antibody\u2019s binding site. Antibody IgG is the smallest, comprised of just a single large receptor that pairs up with the virus it aims to degrade. Antibody IgA consists of a mini adapter and two large receptors, while antibody IgM consists of a small adapter and five large receptors (Figure 1(a, b)). IgA controls a variety of protective functions in the body, while IgM antibodies are the first to be made when a virus enters the body. These receptors pair up with the small receptors, or epitopes, on the viruses\u2019 antigens, which are formed by assembling a large adapter with small receptors (Figure 1(b, c)). This binding occurs through the use of magnets inserted into the 3D prints, enabling the students to depict an analogy of the antibodies moving towards the viruses and inducing their immune response through epitope binding. The kit also contained models of a type of virus called bacteriophage T4, another example of what the antibodies can degrade (Figure 1(d)). Ultimaker Cura software was used to resemble the modelling of the antibody-antigen binding for all three antibodies and the two different viruses (Figure 2). To allow for a more accurate visualisation of the 3D prints, the Ultimaker Cura software was also used to measure the dimensions of the antibodies and virus models [Supplementary Material Figure 5]. The students were given freedom in handling the 3D models as three researchers walked around observing the groups. The researchers aided the students in connecting the pieces to represent the antigen-antibody binding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure5-1.png",
        "caption": "Fig. 5. Forward kinematic singularities: (a) first configuration; (b) second configuration.",
        "texts": [
            " (10), one can derive that the 2UPR-PRU PKM is in the inverse kinematic singularity only if limb 3 is perpendicular to slider 3OB . Fig. 4 shows the configuration. 3.2.2 Forward kinematic singularity When =\u03b7J 0 and \u2260qJ 0, the PKM is in the forward kinematic singular configuration. Eq. (10) allows for the derivation that the 2UPR-PRU PKM is in the forward kinematic singularity only when limb 1 or 2 is co-linear with 1 2A A . Therefore, two forward kinematic singular configurations occur, as shown in Fig. 5. 3.2.3 Combined kinematic singularity Only when =qJ 0 and =\u03b7J 0 are satisfied simultaneously, the PKM is in the combined kinematic singular configuration. These results show that combined kinematic singularity occurs only when limb 1 or 2 is co-linear with 1 2A A , and limb 3 is simultaneously perpendicular to slider 3OB , as shown in Fig. 6. From Figs. 4-6, one can find that a good choice of link parameters can enlarge the workspace region. 4. Dynamic analysis Dynamic modeling of 2UPR-PRU PKM based on the screw theory and principle of virtual work can be used to explain the mapping between the actuated forces and the output trajectory, which is useful for the development of efficient control system in practical applications"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000892_pesgre45664.2020.9070281-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000892_pesgre45664.2020.9070281-Figure1-1.png",
        "caption": "Fig. 1. FEA model of HP fuel pump motor.",
        "texts": [
            " This approach of FEA based modelling of motors possesses several advantages over the high order state-space models made in MATLAB/Simulink environment [16]. First, it considers various effects which impact motor performance, including eddy currents, hysteresis losses and magnetic saturation. Second, it integrates geometrical details, properties of materials and boundary conditions into FEA models. Finally, it facilitates automatic transfer of complete 2D or 3D geometry. The FEA model of HP fuel pump motor is shown in Fig. 1. The current data obtained from the simulated motor model is shown in Fig. 2. Fig. 2(a) shows the current values during the window of 0.1 \u2013 0.7 seconds of the motor operation. Fig. 2(b) shows a zoomed-in version during 0.15 \u2013 0.25 seconds. The model is then validated by comparing the current derived from simulation results with the measurements obtained by actually operating the motor at rated loading conditions. Table I enlists a few simulation parameters of HP fuel pump motor. In a DC motor, the total current is proportional to the difference between the applied voltage and back EMF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure3.9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure3.9-1.png",
        "caption": "Fig. 3.9 Scheme of the high-frequency induction heating system employed in the sintering of metal powders",
        "texts": [
            " The pieces to process may be placed on conveyor belts for their handling in mass production processes. In the case of aluminum alloys, an inert atmosphere is required to prevent the formation of undesirable surface films, such as aluminum oxide [20]. This technique is the most used in the mass production of metal components, as well as in the study of experimental materials at the laboratory scale. 3.5 Sintering of\u00a0Metal Powders 40 However, sintering processes consider additional technologies, where the use of high-frequency plasma or induction systems (HFIS, Fig.\u00a0 3.9) is also explored because of the advantages offered by their high sintering speed rate [21], which is simultaneously carried out with the consolidation process. Unlike conventional sintering processes, which require long periods of heating and cooling, one of the advantages provided by rapid sintering methods is that they provide the amount of heat needed in the area of interest, in the period required by the product in its green stage. This is necessary to carry out the interdiffusion mechanism and the subsequent formation of metallurgical bonds among the metal particles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000185_1.5131964-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000185_1.5131964-Figure3-1.png",
        "caption": "FIGURE 3. The scheme of the print unit of the EBM machine",
        "texts": [],
        "surrounding_texts": [
            "The starting materials were following: welding wire made of Ti-6Al-4V a diameter of 1.2 mm and stainless steel AISI 308LSi with a diameter of 1.2 mm. The chemical composition of the materials is presented respectively in Tables 1 and 2. In 2015 a modular installation for electron-beam fusion of powders and cladding with wire was created and have been constantly modernized at TPU from then on. The setup includes a vacuum chamber with an electron-beam gun with a plasma emitter and modular manipulators, enabling layer-by-layer cladding of powders by electron-beam melting (EBM) or dimensional welding with wire. The appearance of the setup and the structural schemes are shown in Figures 1\u20133. 020097-2 The technical characteristics of the machine are as follows: accelerating voltage is 40 kV, base pressure is 5\u00d710-3 Pa, maximum beam current is 200 mA, minimum current beam diameter is 150 \u03bcm, build area dimensions are 150\u00d7150 mm, power consumption is 6 kW. The electron beam is generated by a source with plasma emitter. A distinctive feature of the source is the use of gas discharge plasma as the electron emitter. Such design provides a number of new operational benefits in comparison with thermal cathode sources. For instance, plasma-emitter sources preserve performance at elevated operating pressure. In addition, they demonstrate long life under the action of metal vapors, including refractory ones, and gas emissions from the melting zone. They are successfully used in traditional electron-beam technologies such as welding, surfacing. The machine is intended for manufacturing parts with complex geometries from metal powders or a metal wire. The included software enables modular exchange and synchronized control of all installation components according to the task using digital G-codes. Initially, the Linux NC system was designed to control turning, milling machines, 3D printers and other mainly machine-building equipment. However, the potential laid by the developers in the system makes it possible to develop a more complex system, for example, an electron-beam installation. During the work, samples of Ti-6Al-4V titanium alloy and AISI 308 steel were printed on an electron-beam 3Dprinter at an accelerating voltage of 30 kV and a beam current of 15 to 20 mA (depending on the distance from the substrate). The input power varied from 450 to 600 watts. The focused beam (150 \u03bcm in diameter) was performing circular scan with 4 mm in diameter. The wire was fed to the beam area, and the sample geometry was achieved by moving the table along three axes. The distance between the tracks (hatch distance) was 4 mm, the layer height was 0.8 mm, the beam was moving in in the horizontal plane in zigzag pattern. The samples shown in Figures 4 and 5 demonstrate the installation capabilities for obtaining both bulk and square samples with wall thickness in one pass. The obtained samples allow studying the structure and microstructure of the alloys depending on the energy, geometric and kinetic parameters of the beam and the geometry of the obtained samples. To assess the porosity of the specimens and to perform mechanical testing, the samples were cut into rectangular specimens shown in Figure 6 by the electroerosion method. 020097-3 Vickers hardness was determined on the transverse sections of the specimens using an Tukon 2500 Automated Knoop/Vickers Hardness Tester (2500-6 modification, Instron, USA) with a load of 1 kg. The indentation time was 10 seconds. \u0425-ray computed tomography (XCT) studies were performed using a scanner OREL-MT developed in TPU. The scanner equipped with XWT-160-TC X-ray tube (X-ray WorX, Germany) and PaxScan-2520V flat panel X-ray detector (Varian, the USA). The scanning parameters were as follows: accelerating voltage \u2013 130 kV, copper filter of 0.2 mm thickness, angle step of 0.3 degrees, 1200 shadow projections, voxel size \u2013 10 um."
        ]
    },
    {
        "image_filename": "designv11_63_0001147_j.jmmm.2020.167119-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001147_j.jmmm.2020.167119-Figure7-1.png",
        "caption": "Fig. 7. winding configuration of 6-phase IM.",
        "texts": [
            " Polyphase slip-ring IM was used for simulation, the machine specifications are given in Table 1. 7. Result and discussions The algorithm of the proposed approach, shown in Fig. 5, is utilized to obtain the magnetic field analysis, mutual inductance and equivalent circuit, and instantaneous torque. The presented results will be performed at 6 phases in the stator and 6 phases in the rotor. The winding distribution is presented in Fig. 6. The machine winding configuration for one pole is illustrated in Fig. 7. By using FEMM software the magnetic field of the polyphase IM can be deduced as shown in Fig. 8. Also, the mesh was applied by the FEMM preprocessor for breaking machine geometry into several small triangular elements. The number of obtained nodes is 145,716 and the number of elements is 291230. The air gap flux density between analytical formulation using Eq. (12) and the FEMM is shown in Fig. 9. In this figure, there is a slight difference between these results. The flux density analysis of a cross-section of the polyphase IM is illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001507_012020-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001507_012020-Figure6-1.png",
        "caption": "Figure 6. Typical loss component of 37.5 kW three phase squirrel cage induction motor",
        "texts": [
            "5 percent of losses may occur due to bearing friction and inefficient cooling fans. Losses in friction and magnetization are related only to the size and design of the motor and not dependent of motor load. The subsequent losses are called as stray load losses, which is around 2 percent of total losses. Partially loaded motors have lower efficiencies because of constant losses such as magnetizing losses, friction losses, and windage losses that will be the larger percentage of total motor power consumption. Figure 6 shows the various components of losses for a 50 hp squirrel cage induction motor as a function of motor load. The overall efficiency of a motor is dictated by the construction materials, mechanical and electrical design. Power sensitive motors use high quality components and use engineered architecture to achieve higher efficiencies. Large copper wire diameter in the stator and more aluminum in the rotor minimize motor resistance losses. Enhanced rotor configuration and optimized air gap between stator and rotor results in stray loss reduction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000089_ev.2019.8892977-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000089_ev.2019.8892977-Figure3-1.png",
        "caption": "Fig. 3. PMVM-SMMRS magnetic flux density map and flux lines distribution",
        "texts": [
            " a) PMVM with surface magnets mounted in rotor and stator poles (PMVM-SMMRS) The outer region is based on 27 negative and 27 positive magnets oriented radial which gives an advantage in magnetizing the magnetic core. Due to the both stators, the rotor is placed between the stators and receives magnetic energy from two sources to interact with PM stuck inside. The core material used for all structures is M270 with 0.35 thickness and the PM are 1.4 T. In this section it is presented the magnetization of the cores and the flux lines passing through inner stator, the inner air gap, the PM in the rotor, the rotor core, the outer air gap, the PMs in the outer stator and the outer stator core. From Figure 3 to Figure 5, a) can be highlighted that the magnetic flux density does not exceed 2.3 T, although PMVM-SMMRS rotor and outer stator is more saturated in comparison with PMVM-BM and PMVM-SMM. The flux lines distribution is represented in Figure 3 to Figure 5, b) where the role of PM can be highlighted. Not all the flux lines produced in the stator secondary poles are closing through the outer stator which results in the influence of the coercive magnetic field produced by the PM in the outer stator. Following the electromagnetic analysis, electromagnetic torque variation in time is represented in Figure 6. PMVMSMMRS develops 108 Nm, an average value of torque with a consistence of 13.35% torque ripples and it is rated at 1696.5 W output power with 95"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure9-1.png",
        "caption": "Fig. 9. Origin measurement error of PCS.",
        "texts": [
            " Therefore, it is usually to control angular error of the pendulum axis to less than 1\u25e6 to meet COG measurement accuracy. The ideal origin point of PCS locates at the ideal pendulum axis if there is no measured error of PCS. However, due to measurement errors, the measured origin point does not fall on the ideal pendulum axis, so the measured pendulum axis will have a positional error with the ideal pendulum axis. If no angular error of the pendulum axis exist, the measured pendulum axis should be parallel to the theoretical pendulum axis, as shown in Fig. 9. The measured and the theoretical origin points are both on the face of the pendulum. Assuming that the measured origin points under two suspended postures of the measured body are Od1m and Od2m as shown in Fig. 9, respectively, and the theoretical origins are Od1 and Od2 respectively. Then the COG can be calculated using Eq. (17): rOr Cm = 1 2 ( rOr Od1m + rOr Od2m + Rm S \u2212 Qm S2 \u2212 1 e1 + Rm \u2212 Qm S S2 \u2212 1 e2 ) (35) where Cm represents the calculated COG when there are measured errors for the origin points of the PCS. The Rm and the Qm are calculated as: Qm = rOd1m Od2m \u22c5e1 = (rOd1 Od2 \u2212 rOd1 Od1m + rOd2 Od2m )\u22c5e1 (36a) Rm = rOd1mOd2m \u22c5e2 = (rOd1Od2 \u2212 rOd1 Od1m + rOd2 Od2m )\u22c5e2 (36b) Assuming no angular measured error of pendulum axis, the vectors e1, e2 and scalar S are unchanged in RCS under two different suspended postures, and the definitions of these three variables in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000792_ica-200622-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000792_ica-200622-Figure6-1.png",
        "caption": "Fig. 6. Angular position of the the master cart\u2019s tilt sensor.",
        "texts": [],
        "surrounding_texts": [
            "The development of the mathematical model that is carried out for the computation of the reference trajectories of the carts when they act with the master or slave roles, denoted as x\u2217MS(t) = [\u03b2\u2217MS(t), \u03b3\u2217MS(t), \u03b1\u2217MS(t)]T and x\u2217SL(t) = [\u03b2\u2217SL(t), \u03b3\u2217SL(t), \u03b1\u2217MS(t)]T , respectively, are introduced. Let us highlight that the behaviours of the carts are different depending on the master and slave roles assigned. These roles are dynamically established according to the instantaneous locations of both carts. Moreover, the angular position of the slave cart must vary when a collision is foreseen with the master cart after the trajectories have been calculated. In this case, the slave cart must follow the opposite path. We have to remark that we will focus our study in the development of the mathematical procedure for obtaining the reference trajectories of the master cart, x\u2217MS(t). According to Fig. 4, the computation of these reference trajectories uses the instantaneous position of the patient in the circular rail, P(t), the direction the person is looking at, vp(t), and the instantaneous configuration of the master cart, x(t) = [\u03b2(t), \u03b3(t), \u03b1(t)]. The procedure for achieving the reference trajectories for the slave cart, x\u2217SL(t), is similar to the model developed for the master cart, but using the instantaneous position of the part of the body to be monitored in the circular rail, Pb(t), instead of the instantaneous position of the patient in the circular rail, P(t), and using the direction to which the monitored part of the patient\u2019s body is pointing, vs(t), instead of the direction in which the patient is facing, vp(t). For that reason, the procedure for obtaining x\u2217SL(t) has not been included. We start the development of the model by obtaining the new position of the master cart in the circular rail, \u03b2\u2217MS(t), from which to observe the direction the person is looking at, is given by one of the two points, denoted as P1 = [x1, y1]T and P2 = [x2, y2]T , of the circular rail that intersects with the imaginary line Rp that results from the projection of the direction the patient is looking at, vp(t), on the XY -plane and whose origin is in his/her current position P(t). Next, we will study which of the two points focuses on the said vector, discarding the other and, finally, we will obtain the position to which the cart will move. This approach allows to geometrically obtain the two cut points of the line passing through the patient\u2019s position, taking the patient\u2019s face as the direction. Figure 5 illustrates a visual description of the angular positions obtained for the master cart. In order to obtain the mathematical description of the line Rp, the patient\u2019s position, P(t), and the direction the patient is looking at, vp(t), are used, obtaining: x\u2212 xp xv = y \u2212 yp yv (4) where x and y are the coordinates of the different points that belong to line Rp. The parametrisation of the circular rail of radius R is defined as: x2 + y2 = R2 (5) Upon operating with Eqs (4) and (5), we obtain: x = xv yv \u00b7 y + ( xp \u2212 yp yv \u00b7 xv ) (6)( xv yv \u00b7 y + ( xp \u2212 yp yv \u00b7 xv ))2 + y2 = R2 (7) By defining m = xv yv and n = xp \u2212 yp yv \u00b7 xv , Eq. (7) can be expressed in the following compact form: (my + n)2 + y2 = R2 (8) Now, after simple algebraic manipulations in Eq. (8), the following result is obtained: (m2 + 1)y2 + 2mny + (n2 \u2212R2) = 0 (9) The values of the resulting cut points P1 = [x1, y1]T and P2 = [x2, y2]T , as depicted in Fig. 5, are computed as follows: y1 = \u22122mn+ \u221a 4m2n2 \u2212 4(m2 + 1)(n2 \u2212R2) 2(m2 + 1) (10) x1 = (y1 \u2212 yp) \u00b7m+ xp y2 = \u22122mn\u2212 \u221a 4m2n2 \u2212 4(m2 + 1)(n2 \u2212R2) 2(m2 + 1) (11) x2 = (y2 \u2212 yp) \u00b7m+ xp In order to discover which point is correct, the parametric equation of line Rp with the value of parameter h = x1\u2212xp xv = y1\u2212yp yv is calculated by substituting the value of the first gotten point in said equation. If h > 0, then point P1 is the cart\u2019s destination point. If h < 0, then the cart must move to point P2. The next step consists in using parameter \u03b2M (t) = arctan ( y1 x1 ) to convert that point of the circumference into an angular position around the circle. Once we have obtained the point to be addressed, the movement of the cart has two possible solutions. The first solution is \u2206\u03b21(t) = \u03b2M (t)\u2212 \u03b2(t), while the second corresponds to the opposite direction. If \u2206\u03b21(t) > 0, which corresponds to a clockwise turn, we have \u2206\u03b22(t) = 2\u03c0+\u03b2M (t)\u2212\u03b2(t). If the turn is counter-clockwise (i.e. \u2206\u03b21(t) < 0), we have \u2206\u03b22(t) = \u22122\u03c0+\u03b2M (t)\u2212\u03b2(t). Bearing in mind all the aforementioned, the new position of the master cart in the circular rail, \u03b2\u2217MS(t), is: \u03b2\u2217MS(t) = min(|\u2206\u03b21(t)|, |\u2206\u03b22(t)|) (12) The following step is to calculate the angular position of the RGB-D sensor\u2019s pan of the master cart, \u03b3\u2217MS(t), which is calculated as follows. The direction of the pan vector of the RGB-D sensor of the master cart, whose angle is \u03b3(t), moves until it occupies the opposite direction to the principal vector of the patient, i.e., \u03b3p(t) + \u03c0, where \u03b3p(t) is defined as follows (see Fig. 5): \u03b3p(t) = arctan ( yv(t)\u2212 yp(t) xv(t)\u2212 xp(t) ) (13) However, as the RGB-D sensor is mounted on a cart rotating around the rail, the turn \u03b2\u2217MS(t) made by the cart needs to be subtracted, thus obtaining: \u2206\u03b31(t) = \u03b3(t)\u2212 (\u03b3p(t) + \u03c0)\u2212 \u03b2\u2217MS(t) (14) \u2206\u03b32(t) = \u03c0 \u2212 \u03b3(t) + \u03b3p(t)\u2212 \u03b2\u2217MS(t) (15) where \u2206\u03b31(t) and \u2206\u03b32(t) represent the two possible turn solutions of the pan movement of the RGB-D sensor based on the turn of the cart (clockwise or counterclockwise) around the rail. After computing Eqs (14) and (15), the angular position of the RGB-D sensor\u2019s pan of the master cart, \u03b3\u2217MS(t), is the following: \u03b3\u2217MS(t) = min(|\u2206\u03b31(t)|, |\u2206\u03b32(t)|) (16) Next, the angular position of the tilt of the RGBD sensor of the master cart, \u03b1\u2217MS(t), is calculated. In this case, the direction of the tilt vector of the RGB-D sensor of the master cart, whose angle is \u03b1(t), moves until it occupies the opposite direction to the principal vector of the patient in the vertical plane defined by angle \u03b3p(t) (see Fig. 5), i.e., \u03b1p(t) + \u03c0, where \u03b1p(t) is defined as follows: \u03b1p(t)=arctan  (R cos [\u03b3p(t)]\u2212 xp(t))2+ (R sin [\u03b3p(t)]\u2212 yp(t))2 zp(t)\u2212 zc (17) where zc denotes the constant value of the RGB-D sensor height. Moreover, the movement of the RGB-D sensor is obtained by considering that the cart is also moving. The objective is, therefore, for \u03b1\u2217MS(t) to occupy the place of the angle opposite to \u03b1p(t), that is, \u03b1p(t) + \u03c0. There are again two rotation possibilities to ensure that the RGB-D sensor reaches the desired point: \u2206\u03b11(t) = \u03b1p(t)\u2212 (\u03c0 + \u03b1(t)) (18) \u2206\u03b12(t) = \u03c0 \u2212 \u03b1p(t) + \u03b1(t) (19) Again, after computing Eqs (18) and (19), the angular position of the tilt of the RGB-D sensor of the master cart, \u03b1\u2217MS(t), is the following: \u03b1\u2217MS(t) = min(|\u2206\u03b11(t)|, |\u2206\u03b12(t)|) (20) Finally, grouping Eqs (12), (16) and (20), the final reference trajectory vector for the master cart, x\u2217MS(t), is shown next: x\u2217MS(t) = \u03b2\u2217MS(t) \u03b3\u2217MS(t) \u03b1\u2217MS(t)  = (21)min(|\u2206\u03b21(t)|, |\u2206\u03b22(t)|) min(|\u2206\u03b31(t)|, |\u2206\u03b32(t)|) min(|\u2206\u03b11(t)|, |\u2206\u03b12(t)|) "
        ]
    },
    {
        "image_filename": "designv11_63_0000834_crc.2019.00024-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000834_crc.2019.00024-Figure4-1.png",
        "caption": "Fig. 4. Wheels of the robot",
        "texts": [
            " Section IV describes a verification experiment and its results, and Section V provides conclusions. The robot used in this research was \u201cTateyama,\u201d a wheeled robot developed in our laboratory (Fig. 3). Table I lists the specifications of the robot. The vehicles are deployed in a forward-and-aft configuration for ascending steps. When the wheelchair and the robot encounter a step, the robot hand grasps the rotary shaft of the push handle of the wheelchair [18]. The robot has three pairs of left and right wheels (Fig. 4). The front pair are casters, and the middle and rear pairs are driving wheels. The arms to which the front and rear wheels are mounted can be deployed and retracted (Fig. 5), and these mechanisms are used in the step climbing process. The robot has manipulators attached to the left and right joints of its upper half; each arm has 5 degrees of freedom (DOFs) and the hand has 1 DOF, for a total of 6 DOFs (Fig. 6). As configured for this study, the length of the upper arm link [from Joint 2 (shoulder) to Joint 4 (elbow)] was l2 and the length of the forearm link (from Joint 4 to Joint 6) was l4C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001848_s38313-020-0306-7-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001848_s38313-020-0306-7-Figure6-1.png",
        "caption": "FIGURE 6 Insulation system of the exhaust port (\u00a9 FEV)",
        "texts": [
            " Initial topology investigations showed that the crankcase with the hollowed bulkheads has a sufficient structural rigidity. The design was adjusted ac - cordingly so that the hollow parts of the crankcase could be used as oil return passages. The design freedoms of AM allow various\u00a0structures to be produced during the process of printing. The main idea of the shown insulated exhaust port was the targeted layout of a heat flux reducing insulation lattice element which can be combined with an air gap, FIGURE 6. This leads to a reduced heat flux from the exhaust gas into the engine coolant. This system re duces\u00a0the heat flux by around 5\u00a0% at rated power. It enables a shorter warm-up period of the exhaust gas aftertreatment system and a higher turbine entry temperature. The distribution of the pillar structure\u00a0was FEA-optimized with the goal of maximizing the exhaust port wall temperatures within the material boundaries, while creating an even wall tem- FIGURE 5 Lubrication system (\u00a9 FEV) MTZ worldwide 12|2020 41 perature distribution across the exhaust port"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002244_012013-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002244_012013-Figure3-1.png",
        "caption": "Figure 3. The scheme of influence of a relative axial gap at the coefficient of friction.",
        "texts": [
            " As the axial clearance of the side wall increases, the friction losses are reduced to a minimum (which is of no practical significance, as it is not really possible to do so) before it increases again. As already mentioned, the effect of the axial gap of the side wall is minimal in turbulent flow, while it is significant in laminar flow. The more the surfaces of the sidewall gap cover get wet, the slower the liquid rotates Pantell's experiments showed that the relative effect of the axial gap in the presence of a casing for a rough disk is the same as for a smooth disk. It is schematically shown in Figure 3. HERVICON+PUMPS 2020 Journal of Physics: Conference Series 1741 (2021) 012013 IOP Publishing doi:10.1088/1742-6596/1741/1/012013 As you can see, the minimum value of the coefficient of friction of the disk is observed at point 2 at the width of the axial gap . With a further decrease in the axial gap (sections 1\u20132), a significant increase in the coefficient of friction is observed. This can be explained by the fact that when applying the boundary layers, the velocity gradient in the fluid flow increases sharply"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003473_s41315-021-00184-1-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003473_s41315-021-00184-1-Figure2-1.png",
        "caption": "Fig. 2 Coordinate frames of the omnidirectional mobile manipulator",
        "texts": [
            ", 1 and 2 ), which are respectively driven by one DC motor. Compared with the serial manipulators, the actuators of parallelogram manipulators are fixed to the base. Parallelogram manipulators have the ability to perform high speed and acceleration motion. Therefore, they are suitable for high velocity gripping and handling of light and small materials (Li and Xu 2007). However, the parallelogram structure also brings complexities in the dynamics modeling. 1 3 The coordinate frames are shown in Fig.\u00a02: the world coordinate frame {W} fixed on the ground and the moving coordinate frame {M} fixed on the robot geometric center. The nomenclature is defined in Table\u00a01. The mechanical structure of robot prototype is designed by the software of Solidworks, and the material property is known. Thus the parameters used in experiments can be roughly estimated through Solidworks. Motor parameters can be obtained from product manual. In addition, it is assumed that no slippage exists between the wheel and the motion surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001788_ecce44975.2020.9235382-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001788_ecce44975.2020.9235382-Figure1-1.png",
        "caption": "Fig. 1: Machines Geometries, (a) SPMIC machine, (b) SyncRel machine, (c) SPMIL machine",
        "texts": [
            " In faulty conditions, the air-gap magnetic field distribution is unbalanced. This asymmetry may be caused either by mechanical faults, e.g. static or dynamic eccentricity; or by electric faults, e.g. fault in a stator phase. Hence, the optimal choice of the machine type shall be based on assessment of radial forces in faulty conditions. Here, the proposed dual-rotor ironless SPM machine is compared with two traditional iron core machines in terms of radial force during faulty conditions and of mutual coupling between the stator phases. Fig. 1 shows the geometries of the three synchronous machines which were designed fixing the stator outer diameter at De = 150 mm and with the same rated output torque TN = 3 Nm at \u03c9N = 9000 rpm. Tables I-II summarize the main characteristics and performance of the reference machines. Radial forces were estimated for the three reference machine in healthy and faulty conditions by 2D FEA using FEMM software. In the following the three reference machines are referred to as SPMIC (iron core surface permanent magnet machine), SyncRel (Synchronous reluctance machine), SPMIL (MechSTOR ironless dual rotor permanent magnet machine)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000686_mepcon47431.2019.9008048-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000686_mepcon47431.2019.9008048-Figure8-1.png",
        "caption": "Fig. 8. Solid model of removed stator.",
        "texts": [
            " LOCKED ROTOR ANALYSIS Locked rotor test (LRT) is performed to obtain rotor resistance and leakage inductance. LRT is simulated by solving the model with zero rotor speed. Magnetizing inductance is assumed to be constant due to the fact that under this test the magnetic circuit is linear since the input voltage is Lls=3 p iph (1) Lls= Wf iph 2 (2) Lm= 3 ph Inl -Lls (3) Lm= Wf iph 2 -Lls (4) # Authorized licensed use limited to: University of Newcastle. Downloaded on June 01,2020 at 15:54:20 UTC from IEEE Xplore. Restrictions apply. low (reduced voltage). Rotor circuit parameters depicted in Fig.8, are determined by LRT by exciting stator winding with full load current and keeping the rotor at zero speed. Therefore, stator and rotor frequencies are equal. Rotor equivalent circuit impedance (Req and Leq) are calculated from rotor copper losses as well as energy stored in air gap as by Eq.5. , and Eq.6, respectively. After determining rotor equivalent impedance, the actual rotor parameters referred to stator are obtained using Eq.7 and Eq.8. Analytical determination of rotor parameters has also been obtained as per [14] for the sake of validation. Results of both techniques are compared in Table 2. In this proposed test, stator lamination\u2019s material is replaced by a material with a very low relative permeability as shown in Fig.8. Consequently, no useful flux is produced, and the flux produced pertains to leakage flux only. The concept is similar to removed rotor method described before. Fig. 9. Flux distribution with removed stator. Fig.9 shows rotor flux distribution with RST. The flux is confined to rotor laminations only and no flux penetrates the stator therefore this flux can be considered as leakage flux. Rotor leakage inductance can be determined as per Eq.9, where Ibar is the peak rotor current obtained from full load performance, and k is the transformation factor (described in detail in [14])"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002010_j.tws.2020.107334-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002010_j.tws.2020.107334-Figure18-1.png",
        "caption": "Fig. 18. LMS model of the cylindrical shell.",
        "texts": [
            " The modal analysis and processing part includes an LMS TEST. Lab 14a client software. The structure diagram of the test system is shown in Fig. 17. In order to reflect the basic shape and characteristics of the cylindrical shell, thirty points are uniformly calibrated on the outer surface of the cylindrical shell. The thirty points divide the cylindrical shell into five sections along the circumferential and the axial directions, respectively, and the corresponding structural model is established in LMS TEST. Lab as shown in Fig. 18. In the experiment, a force hammer is used to knock and excite the fixed excitation points on the cylindrical shell. Six acceleration sensors are pasted along the axial calibration points to obtain the vibration information of the system. Then the acceleration sensors are moved along the circumferential direction to obtain the modal characteristics of the whole cylindrical shell. Fig. 19 shows the layout of the whole active stiffness control test system. The modal characteristics of the cylindrical shell under each sensor layout are measured ten times and averaged"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002511_012004-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002511_012004-Figure11-1.png",
        "caption": "Figure 11. Distribution of maximum stress at point of contact",
        "texts": [
            " Currently, with such modern tools as CAD/CAE software with modules design of machine elements, we can quickly find out these effects of gear module m and pressure angle \u03b1 on contact stress. RCMEManuE 2020 IOP Conf. Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10.1088/1757-899X/1109/1/012004 First, we respectively compare the results calculated with the analytical formula according to ISO and with FEA (ANSYS software). When simulated on ANSYS, the maximum value of contact stress is equal to 371.84 N/m2 (Figure 11). Then, another calculation is made with the analytical formula according to ISO 6336 using Equation (1) is 381,142 N/mm2 (MPa). Compare the % difference between the two results: 383,142 371,84 %ERROR 2,9% 383,142    (5) There is no significant difference between the values obtained when calculations are done respectively by ISO 6336 standard and by FEA on ANSYS software. The stress determined when the researcher conducts calculation using ISO 6363: 1996 is slightly greater. However, according to [6], an error of less than 4% is acceptable; this shows that the simulation performed in ANSYS is compliant with ISO 6336"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002563_09596518211003400-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002563_09596518211003400-Figure1-1.png",
        "caption": "Figure 1. Elastic HFV in the first deflected mode.",
        "texts": [
            " The simulation results for the IGC are given in section \u2018\u2018Simulation results.\u2019\u2019 Finally, a concise conclusion is drawn in the last section. In this section, the elastic HFV dynamic modeling is first established. Then, HFV-target relative motion kinematics is described to form the overall IGC dynamic model. It is generally known that aerodynamic forces, gravity, and engine thrust affect the flexible HFV during the flight process, and lateral vibration due to HFV elasticity can be considered as one of the most considerable effect factors of aerodynamic force. As seen in Figure 1, the elastic HFV can be supposed as a simple Euler\u2013 Bernoulli beam with lateral vibration. The elastic deformation25 in 3D space at a position (x, y, z) in accordance with Figure 1 can be defined as ~d and then it may be written as ~d= X\u2018 i=1 fi(x, y, z)hi \u00f01\u00de which can be described based on generalized coordinate of the elastic HFV and the mode shape related to ith generalized coordinate. Also, it must be noted that the elastic deflection can be modeled using the finite number of modes. Moreover, the crucial factors of mode election for accurate modeling can be regarded as a significant issue. Commonly one needs to select modes that play a more important role in flexible structure excitation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001884_1350650120975519-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001884_1350650120975519-Figure4-1.png",
        "caption": "Figure 4. Finite element model of spherical plain bearings. (a) Finite element model. (b) Loading area on the dabber (marked by red).",
        "texts": [
            " Therefore, a constant weight x is intro- duced into the calculation of the strain to describe the weight of the above effects. The compressive strain e22 of the outer ring in y direction is assumed equal to that of the inner and thus takes the form e22 \u00bc x y z\u00f0 \u00de t \u00fe y z \u00bc B=2\u00f0 \u00de t y z \u00bc B=2\u00f0 \u00de t (8) The compressive stress r22 on the outer ring can then be calculated based on a linear assumption r22 \u00bc Ee22 (9) In order to simulate the practical working conditions of joint bearings,25,29 a dabber is inserted into the inner ring while the outer ring is surrounded by a fixed base, as shown in Figure 4(a). A total of three surface-to-surface contacts are defined in this model, i.e., the contact between the outer ring and the fixed base, contact between the outer and inner spherical surfaces, and contact between the inner ring and the dabber. The friction coefficient l between the inner and the outer rings varies from 0.01 to 0.3 for studying the effects of friction on the contact pressure, while those of the other two contacts are fixed as 0.3 to avoid any possible slipping or rotation. All displacements of the base\u2019s side faces are prohibited to restrain the movement of the whole system with no constraint imposed on the end faces of the outer ring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000162_012003-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000162_012003-Figure9-1.png",
        "caption": "Figure 9. Printed control arm (a) before, (b) during and (c) after the test.",
        "texts": [
            "1088/1742-6596/1386/1/012003 The optimized geometry was manufactured by AM in nylon reinforced with fiberglass, using the parameters in Table 2. The printed parts in which the control arm was divided are shown in Figure 8. The printings were satisfactory although the dimensions were not accurate, especially those of the holes. It was also observed that in the areas where support material had to be removed, the quality of the surface was rough. Table 3 presents the details of each of these parts. The final printed control arm (after joining the two parts and introduce the brass inserts) is presented in Figure 9(a). The weight of this final part was 216 g, which is significantly lower compared to the 835 g of the steel control arm. The assembled suspension of the buggy car with the printed part can be seen in Figure 9(b). The control arm was mounted easily in the suspension, its dimensions were adequate for normal operation and to keep the wheel in vertical position. No failure was observed in the part. Figure 9(c) shows the control arm after the test. In the present work, the topology optimization of a lower control arm of a buggy car produced by AM was performed. First, an FEA of the lower control arm of the suspension system was done to obtain the stress and displacement fields. For this purpose, the forces on the lower arm were calculated using a kinematic model. The resulting loading conditions were low, the maximum force was near 83 N and, therefore, there were also low stresses, close to 20 MPa. These results showed that it was feasible to optimise the shape of the arm to save material and lighten the part"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001623_icra40945.2020.9196598-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001623_icra40945.2020.9196598-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of the CBR descending a slope and the related dynamic parmeters.",
        "texts": [
            " Its motion is characterised, on determinate slopes, by a stable gait in the sagittal plane, without actuation and under the sole action of gravity. The derived mathematical model only describes the swing motion of the nonsupporting leg before the impact with the ground. Afterward, the supporting and nonsupporting leg are swapped. Following [11], an underactuated compass-like biped robot (UCBR) is here addressed by supposing the actuation applied to the ankle of the supporting leg. Referring to Fig. 1, q1 and q2 are both measured with respect to the vertical. The angle q1 is always referred to the support leg, while q2 is always referred to the nonsupport one. Therefore, these angles are not associated to a physical leg, but they are referred to the action played by the leg during the gait. The inertia matrix of the UCBR is given by M(q) = [ (mH + m)l2 + ma2 \u2212mlbcos(q1 \u2212 q2) \u2212mlbcos(q1 \u2212 q2) mb2 ] , (10) with mH > 0 the hip mass, m > 0 the leg mass, a > 0 the distance between the foot and the leg mass, b> 0 the distance between the leg mass and the hip mass and l = a+ b"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure3-1.png",
        "caption": "Fig. 3. Principle of inertia tensor measurement.",
        "texts": [
            " Relationship between the five coordinate systems introduced above are shown in Fig. 2. Assuming the measured body swings a small angle about pendulum axis, the motion equation of the torsional pendulum system is [21]: J\u03c8\u0308 + mgR2 l \u03c8 = 0 (1) where J represents the MOI of the torsional pendulum system about the pendulum axis, \u03a8 is angular displacement of pendulum, R is pendulum radius, m is total mass of the torsional pendulum system including mass of the measured body, the pendulum and the wire, g is gravity acceleration and l represents wire length, as shown in Fig. 3. Eq. (1) is the free vibration equation of a one-degree-of-freedom system. The natural frequency of the system, \u03c9, is equal to the square root of the ratio of constant term to quadratic term coefficient. So the oscillation period can be calculated as: T = 2\u03c0 \u03c9 = 2\u03c0 \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 Jl mgR2 \u221a (2) The MOI of the torsional pendulum system along the pendulum axis is [22]: J = mgR2T2 4\u03c02l (3) When no measured body is suspended on the pendulum, the MOI of the pendulum and the wire about pendulum axis can be calculated using Eq",
            " The error analysis methods for inertia tensor in this chapter can also be applied to traditional TTP for the reason that the measuring principle of these two TTP for inertia tensor measurement is the same [8,9]. Some simplifications are assumed in the process of deriving motion equations for torsional system along pendulum axis and these simplifications may cause model errors. In derivation of Eq. (3), two approximations are assumed: sin\u03c8 \u2248 \u03c8 and sin\u03b6 \u2248 \u03b6. The errors caused by these two approximations will be increased if the two angles become larger. The relation between \u03b6 and \u03c8 is obtained from Fig. 3: \u03b6 = R\u03c8 l (46) In the design of a TTP, it is guaranteed that R\u226al, so \u03b6\u226a\u03c8 . Therefore, model errors caused by the assumption sin\u03b6 \u2248 \u03b6 can be ignored. The relative error of approximation sin\u03c8 \u2248 \u03c8 is: T. Li et al. Measurement 174 (2021) 108956 Em = \u03c8 \u2212 sin\u03c8 sin\u03c8 (47) Relation between relative error Em and initial angular displacement \u03c8 is shown in Table 8: It is known from Table 8 that if the initial angular displacement is less than 5\u25e6, the relative error caused by model linearization is small enough to be neglected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001586_978-981-15-5463-6_82-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001586_978-981-15-5463-6_82-Figure2-1.png",
        "caption": "Fig. 2 Faulty gearset for experimental studies",
        "texts": [],
        "surrounding_texts": [
            "This experimental work was carried out on machinery vibration simulator machine. This machine includes a single-stage bevel gearbox attached to a three-phase induction motor, with a load system. The driver gear has 27 teeth and driven gear has 18 teeth. The experimental setup also comprises a computer-aided DAQ. During the experimental investigation, the characteristics of the vibration signal were evaluated and analyzed in three different conditions of gearbox, i.e., healthy tooth gearbox, broken tooth gearbox, and missing tooth gearbox under the speed of 25 Hz with three specific loading conditions. Figures 1 and 2 show the experimental setup and faulty gearset. Total 3 samples were taken for experimental investigation where, 1 sample for the healthy condition, 1 sample for missing condition, and 1 sample for the chipped condition were collected for 30 s under the three different load conditions (0, 2, and 4 lb) of the gearbox with the speed of 15, 25, and 35 Hz. Figures 3 and 4 show the proposed fault diagnosis method and the raw vibration response for three different cases of the gearbox."
        ]
    },
    {
        "image_filename": "designv11_63_0002263_s00170-020-06477-2-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002263_s00170-020-06477-2-Figure3-1.png",
        "caption": "Fig. 3 Division of cutting areas",
        "texts": [
            " In addition to the large end and small end face of the gear that will have an impact on the shape of the cutting area, the position of the meshing line between the cutter and workpiece on the tooth surface plays a decisive role to the shape of cutting area. In a certain cutting range, the cutting shape does not change; once the meshing line position moves out of the region, the cutting shape will change. Therefore, the convex and concave surfaces of hypoid gear can be divided into five regions according to this rule, as shown in Fig. 3. After the cutting blade cutting into the wheel blank, it intersects with the outer circle of one end face of the wheel blank firstly. At this time, the cutting area is not on the tooth surface. The convex and concave surfaces need to be developed for a period of time before the meshing line can be generated on the tooth surface. The cutting area within this time range is defined as the first stage cutting area. With development of generating, the meshing line continues to move until its two ends intersect with the tooth topland and tooth root respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000555_j.jmapro.2020.01.054-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000555_j.jmapro.2020.01.054-Figure17-1.png",
        "caption": "Fig. 17. Longitudinal residual stress distribution of the components fabricated by AM based on differently shaped substrates in (a) GR-3, (b) PL, and (c) BO-3 schemes.",
        "texts": [
            " Therefore, when the component was cooled to room temperature, tensile stress was generated in the AM part. The material near the bonding zone has the largest constraint on the AM part; hence, it has the highest tensile stress. Finally, a certain range of a compressive stress zone appears around the tensile stress zone to attain equilibrium. However, there are significant differences in the residual stress magnitude and distribution of the components fabricated by AM based on differently shaped substrates. As shown in Fig. 17(a), the maximum longitudinal tensile stress in GR-3 is at the junction of the two sides of the groove and AM part, and is the highest in all models. The reason is that compared with the other two models or other parts in GR-3, additive materials at the bottom of the groove were more strengthened due to more thermal cycles. The model in PL has no AM part inside a 3mm deep groove; hence, its longitudinal high tensile stress zones appear near the height of Z=29mm as the height of the first deposited layer moved up",
            " As the height increases, the compressive stresses occur to attain equilibrium and they are converted into tensile stresses above the bonding zone between the substrate and AM part. The tensile stress of each model reaches a peak at the interface between the substrate and AM part, whose heights are 24mm, 27mm, 30mm, 33mm, and 36mm, respectively. The difference in these residual stress curves is that the tensile and compressive stress zones of different models are distributed differently because their interfaces are inconsistent in height, as also shown in Fig. 17. In addition, although the peak longitudinal stress of the model in PL is minimal among all schemes, their differences are small. A double-peak longitudinal stress distribution is both found in the two models of BO-6 and BO-3, and their ranges of high tensile stress zones are the largest in the five schemes, indicating that the boss structure can enlarge the longitudinal high stress zone. The tensile stress in the AM part of each model is much lower than the peak tensile stress near the bonding zone, indicating that defects are more likely to occur near the interface between the substrate/AM part"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002066_tie.2020.3045591-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002066_tie.2020.3045591-Figure8-1.png",
        "caption": "Fig. 8. Deviation of magnetic saliency from the geometric symmetry pole due to cross-saturation effect (FEA)",
        "texts": [
            " During the drive, effects of several non-ideal factors, such as non-ideal inductance, inverter and current transducer, show up together in measured current responses. These non-ideal factors deform measured current responses and thereby decrease the performance of self-sensing estimations. One of the main non-ideal factors, the cross-saturation effect, is studied as an example to explain the deformation effect on rotor position estimation. In IPMSMs, the crosssaturation effect occurs due to a mutual influence between the flux along d- and q-axes in the same nonlinear magnetic iron. Fig. 8 shows the cross-saturation effect under different load conditions by finite element analysis (FEA). The magnetic saliency aligns with the geometric pole of the rotor d-axis in the case that the cross-saturation effect can be neglected. When the machine operates under no load, the magnetic saliency aligns with the geometric pole of the rotor, as shown in Fig. 8(a). When the magnetic state changes, such as the machine operates under a rated load, the cross-saturation can not be ignored, and the magnetic saliency rotates to ds-axis accordingly, as shown in Fig. 8(b). The self-sensing estimation is consequently subject to an error \u2206\u03b8. Based on Fig. 5, the estimation error of auxiliary angle \u03b3\u0303 has the same modulus of the estimation error of rotor position \u03b8\u0303. Considering the crosssaturation effect only, the location of the deformed point is rotated a certain angle 2\u2206\u03b3 from the primitive point along the circle by the cross-saturation effect, as shown in Fig. 9. By using the rotation operation in transformations T\u2217j , the deformed point G can be shifted to transformed point K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000422_lra.2019.2961302-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000422_lra.2019.2961302-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_63_0000124_012164-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000124_012164-Figure6-1.png",
        "caption": "Figure 6. Safety factor",
        "texts": [],
        "surrounding_texts": [
            "Frame with a maximum load of 1000 N maximum, Safety factors 15 ul and minimum 12.28 ul. The following is an illustration of Safety factors In tables 3 and 4, Figure 11 accepts the greater the force applied (axial loading), the greater the von mises stress, the greater the deflection so that the safety factor gets smaller. Based on the results of simulation analysis using the inventor Autodesk software, the values ranging from loading to maximum 1000 N, amounting to 16.32 MPa while the Results Strength of 207.14 Mpa."
        ]
    },
    {
        "image_filename": "designv11_63_0001707_0309324720958257-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001707_0309324720958257-Figure3-1.png",
        "caption": "Figure 3. Load analysis diagram of straight rope strand.",
        "texts": [
            " Compared to the results of the tension\u2013torsion test of the braided wire rope, the reliability of the mechanical model is verified, which lays a foundation for the friction and wear, failure mechanism analysis, and life prediction of this type of braided wire rope. Mechanical response of wires under tension\u2013torsion load Analysis of force state under tensile load of straight strands The rope strand is a component unit of the wire rope. The analysis of the force state of the straight strand is the basis for establishing the mechanical model of the wire rope. As shown in Figure 3, the stress expression of the core wire and side wire of each layer can be obtained by analyzing the stress state of the straight multi-layer lay rope strand under the tensile load, Ftat. 17 Ftat=F0 +Fn +Fw \u00f01\u00de F0 =T0 =EspR 2 s e0 \u00f02\u00de Fn = n Tn sina 0 n +N 0 n cosa 0 n \u00f03\u00de Fw =m Tw sina 0 w +N 0 w cosa 0 w \u00f04\u00de Where Ftat is the total axial tension of the rope strand, F0 is the axial tension of the core wire of the rope strand, Fn is the axial force along the strand of n inner layer side wires, Fw is the axial force along the strand of m outer layer side wires, T0 is the axial tension of the core wire, e0 is the axial strain of the core wire, Tn is the axial tension of the inner side wire, Tw is the axial tension of the outer side wire, a 0 n, a 0 w are the helix angles of the inner and outer side wires after deformation, respectively, N 0 n, N 0 w are the shear forces in the normal direction of the inner and outer layer side wires, respectively, and m, n are the number of inner and outer layer side wires of the rope strand"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002605_s11071-021-06365-8-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002605_s11071-021-06365-8-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of LOS guidance method",
        "texts": [
            "1 Design of LOS guidance According to the decoupling design of the robotic penguin, the 3D PF problem can be decomposed into planar PF control and depth control, in which the planar PF control can be transformed into yaw angle control. Thus, the yaw angle error and Z-direction error are paid more attention to compensate. From the desired ZPE manifold (3), we can get cross-track error and along-track errorwhich calculated by[ s e ] = [ cos(\u03c8d) sin(\u03c8d) \u2212 sin(\u03c8d) cos(\u03c8d) ] [ xd \u2212 x yd \u2212 y ] . (5) Thus, the target state is expressed as qt = qd , \u03c8t = \u03c8d + atan 2(e, s), rt = \u03c8\u0307t , (6) and the supplementary diagram is described in Fig. 2. 3.2 Design of LESO To tackle the dilemma that real-time velocity is difficult to measure, the LESO is designed \u0307 = [ A C 04\u00d78 04\u00d74 ] + [ B 04\u00d73 ] \u03c4 \u2212 \u03b4, (7) where \u03b4 = [ x\u0302 \u2212 x, y\u0302 \u2212 y, z\u0302 \u2212 z, \u03c8\u0302 \u2212 \u03c8 ]T , =[ q\u0302, \u03c4\u0302 sum ]T , C = [04\u00d74, I4\u00d74]T , = [ 3w0 I4\u00d74, 3w2 0A1, w 3 0A1 ]T , \u03c4\u0302 sum = [\u03c4\u03021, \u03c4\u03022, \u03c4\u03023, \u03c4\u03024]T denotes the sum of model uncertainties and disturbances, and w0 represents the bandwidth of the observer. Specially, hereinafter, q\u0304 = [x, y, z, \u03c8, u\u0302, v\u0302, w\u0302, r\u0302 ]T . 3.3 LMPC scheme The linearized state equation can be obtained by Taylor expansion acting on any desired path point (qt , \u03c4 t ) and leaving the first-order terms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002194_s12206-020-1201-5-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002194_s12206-020-1201-5-Figure6-1.png",
        "caption": "Fig. 6. Schematic map of planetary gearbox structure.",
        "texts": [
            " For ease of description, the planetary gearbox degradation experiment and condition monitoring data collection methods are introduced first. Planetary gearbox degradation data is derived from a lifecycle degradation experiment. The planetary gearbox test rig is shown in Fig. 5. The test rig consists of drive motor, magnetic powder brake, speed and torque sensor, test gearbox and other main parts. The test gearbox is a single-stage NGW-11 planetary gearbox with a transmission ratio of 12.5, the specific structural parameters of test gearbox are shown in Fig. 6. During the experiment, four vibration acceleration sensors are arranged in the planetary gearbox case, and the specific mounted location of sensors are shown in Fig. 7. Sensors 1# and 2# are installed at the input shaft end, sensor 1# is horizontal, sensor 2# is axial direction, sensor 3# is installed at the top of the gearbox, sensor 4# is installed at the output shaft end and is axial direction, and all sensors are installed near the ring gear. In the life-cycle degradation experiment process, the input shaft speed of planetary gearbox is about 1000 rpm, and the load current of the magnetic powder brake is 1 A (about 340 Nm)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000764_bulef48056.2019.9030768-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000764_bulef48056.2019.9030768-Figure1-1.png",
        "caption": "Fig. 1. Automated Roll-to-Roll fluidic self-assembly process. (A) Overview illustrating component dispensing, fluidic self-assembly based on surface tension, and recycling of excess components. (B) Principle of jet pump. (C) Illustration of surface-tension driven self-assembly using molten-solderbased-receptor.",
        "texts": [
            " In addition, we use a novel lamination approach in the realization of flexible solid state lighting modules incorporating distributed inorganic light emitting diodes (LEDs) using the assembly system. Since this application requires not only self-assembly, but also the formation of multilayer interconnects, the lamination approach is crucial. 1057-7157 \u00a9 2015 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. Fig. 1 illustrates the layout of automated roll-to-roll fluidic self-assembly system. The system contains two units: (i) roll-to-roll assembly unit (shaded in gray) involving motor, rollers, customized agitator, and polyimide web to regulate operation parameters such as web moving speed, web angle, agitation frequency and amplitude. (ii) component recycling and dispensing unit (shaded in blue) containing diaphragm pump, jet pump and dispensing nozzle to reintroduce and gently dispense unassembled components",
            " The system operation includes four steps: (i) transporting the components to the assembly unit; a jet pump delivers the unassembled components upward into a narrow fluid channel (5 mm inner diameter tubing). Originally we installed a mechanical pump between the bottom of the chamber and the dispensing head to circulate the components, however, this lead to mechanical damage to the components and the pump. To prevent this damage, we use an indirect circulation approach coupling a mechanical pump (QV variable speed pump, Fluid metering, Inc., NY) with the customized jet pump as shown in Fig. 1(B). The jet pump requires a smaller diameter nozzle (1.48 mm2) to accelerate the carrier fluid into a desired direction. Typical velocity of carrier fluid in the upward fluid channel (20 mm2 tubing) is 25 cm/s to lift up the dense silicon components (density of 2.33 g/cm) to the dispensing head. (ii) dispensing the transported components on the substrate; the components are dispensed on top of the substrate by gravity. Gentle introduction of components to the receptor is important to reduce the effect of liquid flow which induces additional drag. The dispensing head is located 5 cm away from the polyimide web (50 \u03bcm thick, 5 cm wide) pointing not directly at the web. Since the components leave the narrow dispensing head and meet large volume of liquid, the velocity decreases and the components fall downward following a vertical path to the web by gravity. (iii) assembling the components on the substrate; the dispensed components are self-assembled on the advancing substrate. Fig. 1(C) shows the detailed attachment process. The substrates contain multiple pre-defined solder-coated receptors which are formed using a customized dip-coating of a low melting point solder on copper clad polyimide (50 \u03bcm thick, Pyralux LF series, DuPont, NC) films patterned using a microfabrication technique. Procedure to fabricate the substrate is described in the experimental section. The self-assembly process is accomplished in water at 80 \u00b0C where the solder (Indalloy #117, MP. 47 \u00b0C, Indium Corp"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002806_tia.2021.3079169-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002806_tia.2021.3079169-Figure13-1.png",
        "caption": "Fig. 13. Prototype stator, rotor shaft and assembled machine with fan blade.",
        "texts": [
            " In contrast, it is impossible to pass through the critical speed in case of G 2.5 because the radial displacement is rather high even when the acceleration is improved. Therefore, Fig. 12 is a valuable design guideline to pass through the critical speed in one-axis actively positioned bearingless motors. IV. PROTOTYPE MACHINE AND MEASUREMENT SYSTEM Figures 13 (a), (b), (c) and (d) show the prototype stator, the stator passive magnetic bearing, the rotor shaft and the assembled machine with the fan blade. In Fig. 13(a), the stator Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 16,2021 at 06:27:18 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (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. core material is laminated silicon steels of 35H360. The number of conductors is 90 turns per slot in the center stator core. There are linear hall ICs in between the stator teeth to detect the rotor rotational angular position. In Fig. 13(b), two ring permanent magnets are installed in a case of nonmagnetic material. The permanent magnet material is NdFeB of N40SH. In Fig. 13(c), the rotor shaft consists of two sets of eight-pole segment permanent magnets and passive magnetic bearings. The permanent magnet material is the same with the ring permanent magnet. The eight-pole segment permanent magnets are covered with a plastic holder. In the prototype rotor shaft, the rotor mass eccentricity is not compensated. In Fig. 13(d), a fan blade is assembled with the rotor shaft. The fan blade is made of acrylic material. Aluminum tapes are attached on the surface of the fan blade as the sensor target to detect the shaft vibration by laser sensors. The machine parameters without and with the fan blade are shown in Table I. The rotor mass, radial stiffness and tilting stiffness are measured values. On the other hand, inertias and critical speeds are calculated values. Figures 14 (a) and (b) show measurement systems of the rotor shaft vibration without and with the fan blade"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure1-1.png",
        "caption": "Fig. 1. Eccentricities. (a) Static eccentricity. (b) Rotating eccentricity.",
        "texts": [
            " UMF does not occur in symmetrical machines, such as 12- slot/10-pole PM machines whilst it occurs in asymmetric machines, such as 9-slot/10-pole PM machines. In addition, the dominant harmonic order of UMF x/y components is multiple of pole number 2p, where p is the pole pair number. Eccentricity faults, which are usually caused by defective manufacturing or assembly processing [10], also lead to UMF, regardless of machine rotational symmetry [11]-[13]. Eccentricity faults can be classified into static and rotating eccentricities, which are shown in Fig. 1. Static eccentricity mainly adds constant components to UMF x/y components whilst rotating eccentricity adds the 1st order harmonic to UMF x/y components [14]. Another major concern in PM machines is irreversible demagnetization risk. Demagnetization usually happens due to large armature currents, such as short circuit (SC) current, or over-heat of PMs. Practically, certain level of demagnetization is usually allowable to improve working currents and thus power density, or lower PM costs. Therefore, it is also worthwhile to investigate post-demagnetization UMF",
            " 5(a). However, with static or rotating eccentricity, the demagnetization is not rotationally symmetrical and gathers around the PMs facing larger air-gap length. This can be explained by open-circuit and armature field working points, and the influence of static eccentricity on symmetrical STPSPM machine is analyzed as an example. It is defined that positive x direction in the stator reference is 0 mechanical degrees whilst the counter-clockwise direction is the positive direction, as shown in Fig. 1. As aforementioned, the static eccentricity is assumed to happen in the positive x direction. Therefore, for the position 0 mechanical degree is the position with the smallest airgap length whilst the rotor position 180 mechanical degrees is the position with the largest airgap. It can be found that the flux density in the magnetization direction of these two fields decreases when air-gap length increases, as shown in Fig. 14. However, the variation of armature fields is much smaller than open-circuit fields"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure1-1.png",
        "caption": "Fig. 1. Overall diagram of the side-phase anti-bird cage",
        "texts": [],
        "surrounding_texts": [
            "In response to the above-mentioned problems, this paper designs a variable-angle side-phase anti-bird cage. The new side-phase anti-bird cage is aimed at specific types of transmission line towers, taking into account the possible differences between the design drawings of the transmission tower and the reality, and the angle change is used to achieve the attachment of the device to the transmission line tower, ensuring that the protected area of the device will not be allowed birds enter the nest. The shape of the device is a triangular prism composed of five grid surfaces, in which the shapes of the bottom surface and the top surface are isosceles trapezoids, and the sizes of the two are the same, which is convenient for mass production during processing. The back side is a stretchable structure composed of two parts. The sides are two triangular mesh surfaces of the same size. The grid size on all sides of the device is 10mm*10mm. This size can ensure that even the smallest birds cannot enter the nest protected by the device in the area where the device is placed."
        ]
    },
    {
        "image_filename": "designv11_63_0002362_012112-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002362_012112-Figure1-1.png",
        "caption": "Fig. 1 Body coordinate system of four-rotor UAV",
        "texts": [
            " In the reconnaissance mission, when the UAV arrives at the predetermined waypoint to collect images, it needs to adjust the yaw angle several times, which usually takes a long time to adjust to the preset attitude angle. The quadrotor UAV is a strong coupling system with six degrees of freedom, so it is difficult to obtain the transfer function of UAV yaw channel directly. ISPECE 2020 Journal of Physics: Conference Series 1754 (2021) 012112 IOP Publishing doi:10.1088/1742-6596/1754/1/012112 As shown in Figure 1, the UAV body coordinate system is established with the mass center of the quadrotor UAV as the origin ,  ,  They are roll angle, pitch angle and yaw angle of UAV. According to Newton Euler equation[16], 3*3 0 * 0 VmIF mV I I                            (1) In the above formula F For the external force acting on the center of mass of UAV, Is the torque vector of UAV, m For UAV quality, I Is the inertia tensor matrix,V and They are velocity vector and angular velocity vector of UAV. Let the angular velocity of attitude angle adjustment of UAV be  , , T bW p q r In which p q r\u3001 \u3001 They are roll angular velocity, pitch angular velocity and yaw angular velocity in UAV body coordinate system 2 2 ( ) ( ) ( ) ( ) x x xz z y xz y y x z xz z z xz y z xz M q I pr I I p q I M r I p I pq I I qrI                            (2) Among them, xM , yM , zM It is the closing external torque in the direction of three axes in the coordinate system of the engine block, xI , yI , zI The three axes of inertia are the coordinates of the body, xzI It's the product of moment of inertia"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure43.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure43.2-1.png",
        "caption": "Fig. 43.2 CAD model and exploded view",
        "texts": [
            " Mechanism was very similar to umbrella in inverted formation. Tightening screws were for adjustment and was a pass through design in upper frame. There were two different size of plates used plate A = 450 mm and plate B = 300 mm. Plate A was used in couples and was attached to lower frame, whereas plate B was connected with plate A at middle and is fixed with upper frame. All the plastic parts were developed using 3D printers, plates used were of mild steel with 2 mm thick, and transparent plastic sheet is 0.15 mm PVC (Fig. 43.2). Broom was extendable from middle. This was most common type of broom used in household and other minor purpose. So this broom was selected. Body of the broom was of thin steel pipe. Bristles of plastic and since broom was for ceiling it was preferred to be flat and wide for larger area of sweep. Cost of the broom is Rs. 150/-. Specification Broom (length = 2 m, diameter = 20 mm, 0.190 kg) Upper frame was the structure that holds the whole frame of the broom and moving parts in alignment. Later was fixed on the rod of the broom at 10 mm below the brush head of the broom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000300_iemcon.2019.8936288-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000300_iemcon.2019.8936288-Figure15-1.png",
        "caption": "Fig 15- System Flow Diagram explaining the entire Web Service Chain",
        "texts": [],
        "surrounding_texts": [
            "\u2022 The system mainly consists of a RF (radio frequency) module controlled by ARDUINO UNO. RF Module is a cheap wireless communication module used for low cost applications. RF Module comprises of a transmitter and a receiver both of which operate at radio frequency range. Usually, the frequency at which these modules communicate will be 433 MHz. \u2022 The RF module mainly transmits the data of the accelerometer which is to be fitted on the hand glove. \u2022 Jumper wires are used to connect the Arduino board and the transmitter module. Fig.16 shows the entire connection diagram of Arduino and RF TX module. [7] \u2022 Then the accelerometer sensor will also be connected with that Arduino. \u2022 The circuit is our main transmitter circuit which is to be fitted on the glove. \u2022 In order to increase its range and also the strength, we have connected an antenna to the transmitter module. Then the final code is uploaded into the Arduino installed on the hand glove. \u2022 After that, the receiver is connected to the main Robot\u2019s Arduino. Fig.17 shows the circuit diagram of Arduino and RX of RF. [7] \u2022 Here, in the code according to the data received from the accelerometer, the robot\u2019s motors are rotated in accordance with that received data. \u2022 The accelerometer\u2019s data is mapped along the X axis to \u2013X axis for forward and backward direction. Along Y axis to \u2013Y axis, movements from left to right direction is mapped. \u2022 There is a certain gesture where the palm is to be held at a horizontal level which indicates the motors to stop their rotation. \u2022 The motor driver is installed with all the necessary connections. The code is uploaded into the Arduino UNO. Finally, the entire setup is complete."
        ]
    },
    {
        "image_filename": "designv11_63_0001226_acpee48638.2020.9136520-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001226_acpee48638.2020.9136520-Figure7-1.png",
        "caption": "Fig. 7. The diagram of analysis for defects position influnence on the electric field distribution",
        "texts": [
            "2 When there are conductivity defects of different lengths in the both strings of the V-type insulators, the maximum electric field of insulators appears on the insulator string with the longest defect length. If the defect lengths of the left and right insulator strings are equal, the electric field of the defect on the left insulator string which is near to the tower body has a larger value due to the influence of the transmission line tower. C. Influence of the Defect Position in the Single Insulator String a) Defect in the Single Insulator String: The diagram of influence analysis for conductive defects positions on the electric field distribution is shown in Fig.7. In Fig.7, the variable S0 is the distance between the defect and the high voltage end of the insulator. The defect is at the floating potential. When the defect length is 0.5m and the defect position parameter S0 is 0.2m, the calculation results of the electric field distribution are shown in Fig. 8. 1581 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on July 27,2020 at 03:12:24 UTC from IEEE Xplore. Restrictions apply. The electric field with the defect at various position parameters are shown in Table V, and the length of the defect is 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001075_iccar49639.2020.9108090-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001075_iccar49639.2020.9108090-Figure3-1.png",
        "caption": "Figure 3. Dynamic model. (a) The model of single-limb stance phase. (b) The model of double-limb stance phase; (c) A separated model of front and",
        "texts": [
            " The results are given as follows, 1 1 1 1 2 1 1 2 1 1 2 1 2 3 4 sin sin sin( ) sin sin( ) sin sin( ) sin( ) 0 T y y s y s t p y s t y s t t P P L P L L P L L P L L L               \u2212    \u2212 +    \u2212 + + + =   \u2212 + + +   \u2212 + + + + + + +      J () 2 1 1 1 2 1 1 2 1 1 2 1 2 3 4 cos cos cos( ) cos cos( ) cos cos( ) cos( ) 0 T x x s x s t p x s t x s t t P P L P L L P L L P L L L                   \u2212    \u2212 \u2212 + =   \u2212 \u2212 +   \u2212 \u2212 + \u2212 + + +      J ()   3 0 0 0 0 0 0p =J () 0 0 0 0 0 0 0 0 0 0 0 0 1 1 1 1 1 1      =      J () The stance phase of human gait can be divided into two stages [10]. When only one leg supports weight is single-limb stance phase, and when both legs support weight is doublelimb stance phase. In different stages, the constraints are different, which induces different dynamics parameters. Therefore, it is necessary to establish dynamic models for different stages of the gait walking cycle. Fig. 3 shows the dynamic models for different stages, and the centroid position and the centroid velocity of each rod can be expressed using a mathematical expression that includes parameters im , i , i , il and id . Then the Lagrange equation is used to establish dynamic model of the five-bar model. It should be noted that the potential energy of the system is independent of the joint angle i . Lagrange equation can be expressed as formula (10), ( ) ( , ) ( )i = + +T D \u03b8 \u03b8 H \u03b8 \u03b8 \u03b8 G \u03b8 () where ( )D \u03b8 is an inertia matrix, which is a 5\u00d75 positive definite symmetric matrix, ( , )H \u03b8 \u03b8 is a 5\u00d75 Coriolis item, ( )G \u03b8 represents the gravity term, which is a 5\u00d71 matrix, and \u03b8 , \u03b8 , \u03b8 and iT are 5\u00d71 matrices representing the joint angle, angular speed, angular acceleration and joint moment in the generalized coordinate system, respectively",
            " The elements of ( )D \u03b8 , ( , )H \u03b8 \u03b8 and ( )G \u03b8 are calculated as follows, cos( ) sin( ) sin ij ij i j ij ij i j j i i i D p H p G g       = \u2212  = \u2212  =  =     =     () 5 2 2 1 2 5 1 ( ) , 3 , 3 ( ) , 3 , 3 i i i K i K i i i i ij j j i K i j K j j j i ji I m d m l i j and i I m d i j p m d l m l l j i and j m d l j i and j p = + = +  + + =    + = =  =  +      =     () 5 1 ( ) , 3 , 3 i i K i K ii i i m d g m l g i g m d g i = +  +  =   =  () where Dij refers to the element in the i-th row and j-th column of ( )D \u03b8 , Hij represent the element in the i-th row and j-th column of ( , )H \u03b8 \u03b8 , and Gi is the element in the i-th row of ( )G \u03b8 . Unlike the single-limb stance phase, when the double-limb are both contact with the ground, a closed loop is formed as shown in Fig. 3(b). During the period of double-limb stance phase, the velocity of the upper extremity is low and it is nearly a constant, so it can be considered as a quasi-static problem. Then the upper torso mass can be divided into two parts, which are both three-degree-of-freedom linkage systems. The front and rear legs can be analyzed separately. It is assumed that the mass of the torso acting on the front leg is 3m and the mass on the rear leg is 3m , as shown in Fig. 3(c). The distance between the front foot and the center of the rod 3 in X0 direction is xq3 and the distance between the rear foot and the center of mass of the rod 3 in X0 direction is xh3. We can get the dynamic mathematical model of the rear leg as follows, cos( ) sin( ) sin ij ij i j ij ij i j j i i i D p H p G g       = \u2212  = \u2212  = () 3 2 2 1 3 1 ( ) , ( ) , i i i i j i j i ij j j i K i j K j ji I m d m l i j p m d l m I I j p = + = +  + + =   = +        () 3 1 ( )i i i j i j i g m d g m l g = + = +  () The dynamic mathematical model of the front leg is similar to the model of the rear leg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003187_s42417-021-00351-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003187_s42417-021-00351-5-Figure1-1.png",
        "caption": "Fig. 1 Free-floating elastic-base space robot system",
        "texts": [
            " Then, the decentralized fault-tolerant controller is proposed to deal with actuator faults and the uncertain dynamics in the slow-vary subsystem, while a PD feedback controller is designed to address the vibration of the elastics base. Subsequently, \u201cNumerical Simulations\u201d presents the comparative simulation results between the hybrid fault-tolerant controller with a conventional hybrid controller in different situations. Finally, the research ends with conclusions and a reference list. The planar structure of free-floating elastic-base space robot system is shown in Fig.\u00a01. The truss between the base and the guide rail is equivalent to a light spring [16]. The system consists of the free-floating elastic base B0 , rigid links B1 and B2 . OXY is the inertial coordinate frame, while oixiyi (i = 0, 1, 2) is the local coordinate frame of Bi; 0 is the attitude angle displacement of the base, while i (i = 1, 2) is the joint angle displacement of the link Bi . Other symbols are defined as follows: Oi The rotation center of Bi (i = 0, 1, 2) mi The mass of Bi(i = 0, 1) M The total mass of the system, i",
            " To further understand the idea of singular perturbation control, the structure diagram of the hybrid controller is given in Fig.\u00a02, where d indicates the desired trajectories of the base attitude and joints. The slow-varying controller can realize the trajectory tracking control of the base attitude and joint against the actuator faults, while the fast-varying controller can realize the vibration suppression of the elastic base. In order to verify the validity of the proposed FTC algorithm, numerical simulations are conducted for the elasticbase space robot system shown in Fig.\u00a01, and the simulation results of the hybrid fault-tolerant controller are compared with those of the hybrid observer-based sliding mode controller (OSMC) proposed in [21]. To facilitate the description, the hybrid fault-tolerant controller and the hybrid observer-based sliding mode controller are defined as FTC algorithm and OSMC algorithm, respectively. The mathematical expression of the observer-based sliding mode controller of slow-varying subsystem is where = \u22121 rr , = \u22121 rr rr ; k1 , k2 , and kf are positive constants; d = [ 0d, 1d, 2d] T are the desired trajectories; \u0302r are the estimations of the angular displacements r"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001440_s00170-020-05863-0-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001440_s00170-020-05863-0-Figure3-1.png",
        "caption": "Fig. 3 Contour plot of temperature distribution results towards the end of substrate",
        "texts": [],
        "surrounding_texts": [
            "The increased speed of the reinforcement particles that displays a proper flow of molten liquid is brought about by the increase in laser power; also noting that an increased laser power causes improvement, standardized distribution, and spheroidization as shown in Figures 3, 4, and 5. As the laser beam moves from one point to another as shown in Figs. 3, 4, and 5, the temperature changes and the microstructures are determined by the laser input, scanning speed, and fast cooling rate. Figure 6 shows the microstructure of the titanium alloy base metal which acts as heat sink. High cooling rates form dendritic structures within the coatings. The major factors that determine the formation of the dendritic structure are the thermal gradients within the substrate during cooling and the cooling rates as shown in Fig. 7. The faster and slower cooling rates cause the spacing of the dendritic arm. Dendrite arm spacing has great influence on the mechanical properties of additive manufactured parts. The dendrite arm spacing is divided into primary and secondary. Crystal structures are modified by the rate of solidification in the microstructure. The distance from the Ti-6AI-4V alloy substrate determines the peak temperature distribution in the molten pool. There is a notable affiliation between the powder feed rate and laser scanning speed. The turbulent Marangoni convection current is the momentum and turbulence that occurs in the melt pool, which in fact is brought about by the blowing of the powders into the melt pool through the argon gas which carries the powders during laser metal deposition [20, 21, 26]. A decrease in the aspect ratio of columnar grains comes into effect by the increase in equiaxed grains as the build height increases. Processing conditions on grain structure with its influence were also noted. The microstructural development can also be affected by other processing parameters such as the powder flow rate as suggested by the results. Therefore, the microstructural development may have been conclusively determined by the input energy density, which could be defined by laser power, scanning speed, powder flow rate, and thermal history. Hence, there is influence on microstructure when any processing parameter affects the energy density and thermal history. The increase in line energies reduces its influence on the molten pool width, as confirmed by the simulations results. Only for various process parameter combinations were the molten pool depth determined by simulations, with constant line energy resulting in near constant melt pool depth. Higher line energies increased the molten pool depth. The temperature history within the material is prevailed by the process, which influences the mechanical properties of the fabricated coatings."
        ]
    },
    {
        "image_filename": "designv11_63_0002739_j.matpr.2021.04.018-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002739_j.matpr.2021.04.018-Figure4-1.png",
        "caption": "Fig. 4. Stress Plot \u2013 Isometric View.",
        "texts": [],
        "surrounding_texts": [
            "The Baseline model which was analyzed gave results well below the fatigue limit and stress margin. This showed that the model has extensive room for weight optimization. The weight of the baseline Steering Knuckle is around 4.8 kg. After reducing the thickness at all zones where the stresses were very well below the margin, the weight has dropped to 4.0 kg, which is closed to around 0.8 kg or 18%. Table 1 and Fig. 5 show the zones where the thickness was decreased and the weight has been reduced Table 2. The optimized model results showed a max average stress of 80 MPa on the Steering Knuckle, whereas a concentrated band of stress of 189 MPa due to stress concentration is also observed. This maximum stress due to stress concentration due to stress singularity is ignored, since such sharp corners do not exist in real life scenario. Detailed Deformation and Stress Plots are shown as in Figs. 6 and 7, respectively. Table 1 Variation of Dimension on Knuckle. Zone Old Thickness New Thickness Zone 1 & 2 11.2 mm 9mm Zone 3 10 mm 8mm Zone 4 20 mm 16 mm Fig. 5. Optimized Model. Fig. 7. Stress Plot for optimized model."
        ]
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure16-1.png",
        "caption": "Fig. 16 Modeling results of various push\u2013pull edits in a row using the proposed method (circles indicate changed parts)",
        "texts": [
            " Siemens NX failed to update the model for all of the three situations, while the proposed method can successfully update the model for all of them. The comparisons in case studies 1\u20135 are sufficient to show that the proposed method outperforms the state of the art in terms of robustness. Nevertheless, only translational push-pulls were used in these case studies, and the push\u2013pull ranges were small. One more case study (Fig.\u00a015) was thus carried out to show that this work also applies to rotational push\u2013pull with a long range in the edit. The case study shown in Fig.\u00a016 further demonstrates effectiveness of the proposed method by including multiple push-pulls in a row and different push\u2013pull types (translational and rotational). Specifically, four push\u2013pull edits were performed in a row, containing both the rotational push\u2013pull type (e.g., the first push\u2013pull operation) and the translational push\u2013pull type (e.g., the third push\u2013pull operation), as well as single-face push\u2013pull operations (e.g., the second push\u2013pull operation) and multiple-face push\u2013pull operations (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure3-1.png",
        "caption": "Fig. 3. 3D model of Permanent Magnet Synchronous Machine.",
        "texts": [
            " This means, instead of doing the simulation of the complete motor, only one of its symmetric components is simulated and then the complete motor behaviour can be extrapolated. Fig. 1 shows the RMxprt model of the studied machine. When the RMxprt model was analysed, it was exported to the two-dimensional and three-dimensional model using the ANSYS Maxwell Software. In this case, the machine has been divided into four equal parts and only one part is represented on 2D and 3D model as shown in Fig. 2 and Fig. 3. The geometrical of 2D and 3D model are depicted in Fig. 2 and Fig. 3. 682 Authorized licensed use limited to: Raytheon Technologies. Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. During the FE simulation of permanent magnet synchronous machine, the meshing is essential process which is done to descretize the geometry developed into small number of parts called cells. A quarter geometry of the PMSM is subdivided into triangular elements (1000 elements). As obvious from Fig. 4 and Fig. 5 both models give relatively the detailed view of the mesh quality in the 2D and 3D model of the PMSM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure6-1.png",
        "caption": "Fig. 6. Middle-Phase modeling diagram of transmission line tower",
        "texts": [
            " The new device takes into account the influence of the tower angle steel on the sealing performance of the device and eliminates the gap caused by the angle steel structure. The new device is still composed of two parts, both of which are triangular prisms composed of grid surfaces. After the two devices are spliced and combined, there will be no gaps left in the joints of the two devices. According to the drawings of the middle-phase structure of the transmission tower, the middle-phase structure is modeled on the 3D software. The modeled 3D diagram is shown in Fig.6. Looking at the middle phase structure from the left side of the hanging point of the cross arm, you will find that there is no protruding angle steel on the left part of the middle phase, and the anti-bird device can fit well on the left side of the hanging point. The design of the device will not be affected by the angle steel. Looking at the structure at the middle phase position from the right side of Fig.6, you will find two V-shaped angles protruding outwards. When the surface of the right device is a complete mesh surface, it will be blocked by the angle steel, and cannot fit with the vertical surface of the angle steel where the hanging point is located, and there must be a gap on the right side. We have already introduced that middle-phase\u2019s anti-bird device is composed of two parts, both of which have a triangular prism shape. The state of the device after installation is shown in the Fig.7. According to the modeling diagram of the middle phase part of the pole tower, we can find that the right side of the middle phase has an impact on the sealing of the device due to the presence of the angle steel structure, which is the cause of the gaps in the existing antibird devices"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure5-1.png",
        "caption": "Fig. 5. The dimensional parameters for calculating HTS-excitation coil strain.",
        "texts": [
            " When the winding tension is over large, the inner layer of the HTS-excitation coil will be excessive bending, causing the degradation of mechanical properties. Therefore, it is of great necessity for the HTS-excitation coil to consider the proper bend radius and try to reduce its bending stress. When the coil is bent, the inside parts shorten and the outside parts extend. Due to the continuity of the deformation, there must be a layer in the middle that neither stretches nor shortens, namely the neutral layer. As shown in Fig. 5, the strain \u03b5 can be defined as \u03b5 = (\u03c1+ y) d\u03b8 \u2212 \u03c1d\u03b8 \u03c1d\u03b8 = y \u03c1 (1) where y is the distance to neutral layer and \u03c1 is the radius of curvature of neutral layer. Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 11:27:00 UTC from IEEE Xplore. Restrictions apply. Assuming that the elastic modulus of the HTS material are equal in compression and tension, according to Hooke\u2019s law, the bending stress F can be expressed as F = E\u03b5 = E y \u03c1 (2) where E is the elastic modulus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000434_sensors43011.2019.8956932-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000434_sensors43011.2019.8956932-Figure5-1.png",
        "caption": "FIGURE 5. (a) Position trajectories of the multi-vehicle system; (b) Velocity trajectories of the multi-vehicle system.",
        "texts": [
            "0367 0.3767 0.2374 0.7099 \u22120.2053 0.2833 0.7099 2.4161 . Solid blue lines in Figs. 5(a) and (b) respectively show the velocity and position trajectories of the multi-vehicle system from t = 0s to t = 40s. Using the star, asterisk, square and diamond to represent the states of each vehicle, the TVF of the multi-vehicle system under switching interaction topologies is illustrated at t = 10s using bold dash-dotted lines, at t = 25s using bold dashed lines and at t = 40s using bold dotted lines. From Fig. 5, we can observe that both the velocities and positions of the multi-vehicle system reach parallel rectangle formations and the parallel rectangles keep rotation. In Fig. 6, the switching signal, formation error, and coupling weights are displayed respectively. Fig. 6(a) shows that the interaction topology G\u03b4(t) of the multi-vehicle system switch every 1s among G1, G2, G3 and G4 randomly. From Fig. 6(b), the formation error of the multi-vehicle system converges to zero which means that the TVF is achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000371_j.prostr.2019.12.041-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000371_j.prostr.2019.12.041-Figure3-1.png",
        "caption": "Figure 3: Coordinate-system to identify orientation of built to load direction",
        "texts": [
            " With respect to the build time when using the Laser Powder Bed Fusion technology a small size flat specimen geometry, Fig. 2, was used. The specimen geometry was optimized and contains an 11mm long homogeneous cross section, which is necessary for the usage of a clip-on strain gauge. Furthermore, the transition zone from the test section to the clamping section enables the printing of the specimens without support structure as long as the built direction is parallel to the loading direction. In order to describe the specimen orientation, the coordination system according to Fig. 3 was used. In case of Z specimens the built direction is parallel to the load direction, whereas in case of X direction the built direction is perpendicular to the load direction. Rainer Wagener et al. / Procedia Structural Integrity 19 (2019) 380\u2013387 381 Available online at www.sciencedirect.com ScienceDirect Structural Integrity Procedia 00 (2019) 000\u2013000 www.elsevier.com/locate/procedia 2452-3216 \u00a9 2019 The Authors. Published by Elsevier B.V. Peer-review under responsibility of the Fatigue Design 2019 Organizers",
            " With respect to the build time when using the Laser Powder Bed Fusion technology a small size flat specimen geometry, Fig. 2, was used. The specimen geometry was optimized and contains an 11mm long homogeneous cross section, which is necessary for the usage of a clip-on strain gauge. Furthermore, the transition zone from the test section to the clamping section enables the printing of the specimens without support structure as long as the built direction is parallel to the loading direction. In order to describe the specimen orientation, the coordination system according to Fig. 3 was used. In case of Z specimens the built direction is parallel to the load direction, whereas in case of X direction the built direction is perpendicular to the load direction. 382 Rainer Wagener et al. / Procedia Structural Integrity 19 (2019) 380\u2013387 Author name / Structural Integrity Procedia 00 (2019) 000\u2013000 3 Due to the different test requirements of strain- and stress-controlled fatigue test two different types of test rigs were used. In order to perform stress controlled fatigue tests with constant amplitudes, a piezo driven test-rig was used, Fig",
            " Due to this reason, the stresses and strains as well as the Young\u2019s modulus have to be interpreted as a resulting stress, strain or stiffness of structure elements. Well knowing that due to stress concentrations and irregularities the local stresses and strains could be higher. Rainer Wagener et al. / Procedia Structural Integrity 19 (2019) 380\u2013387 383 Author name / Structural Integrity Procedia 00 (2019) 000\u2013000 3 Figure 2: Specimen geometry for strain- and stress controlled fatigue tests to investigate influences on the cyclic material behavior related to additive manufacturing Figure 3: Coordinate-system to identify orientation of built to load direction Due to the different test requirements of strain- and stress-controlled fatigue test two different types of test rigs were used. In order to perform stress controlled fatigue tests with constant amplitudes, a piezo driven test-rig was used, Fig. 4. The specimen is fixed in clamping devices, which are directly mounted on the piezo stack actuator on one side and on the load cell on the other side. In this way, the dynamic properties and the high resolution of the piezo stack actuator can directly be employed to fatigue the specimens"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000415_icoiact46704.2019.8938489-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000415_icoiact46704.2019.8938489-Figure6-1.png",
        "caption": "Fig. 6 Electronic Diagrams",
        "texts": [
            " The moving of Segway is compose of two types of hardware such as electronic and structure. Those of them are presented in this session. Arduino MEGA2560 is the main controller which applied to read sensor and generate the control rule. And ESP 8266 is IOT module, applied to send the data command and output response to monitor. The main sensor is gyro sensor which use to measure pitch angle and command moving direction of the Segway. There is two potentiometers to measure direction of the Segway (Turn Left-Right). And input of the weight of driver to adjust controller gain. Fig.6 shows electronic diagrams of the Segway. Structure of the Segway modifies from the old structure. Almost all are similar to the original. The 500W DC motor is replace to the old one. To transmission power to wheel, gear box is applied with ratio 5:30. To control the heading of Segway, the Segway handle is designed to turn left and light. The illumination of implemented Segway shown in Fig.7 VI. EXPERIMENTAL RESULTS After simulation and implement Segway, the control algorithm is test in real hardware"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000377_012003-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000377_012003-Figure3-1.png",
        "caption": "Figure 3. Initial, intermediate and final positions-orientations of the mobile platform while executing motion 1 (a, b, c) and motion 2 (d, e, f)",
        "texts": [
            "1088/1742-6596/1426/1/012003 The lengths L1\u00f7L6 between the centers of the spherical joints of the same kinematical chain, from 1 to 6, are presented in Table 1. International Conference on Applied Sciences Journal of Physics: Conference Series 1426 (2020) 012003 IOP Publishing doi:10.1088/1742-6596/1426/1/012003 The considered motions are combinations of translations along x and z axes and a pitch type rotation around y axis (according to the Oxyz reference system presented in Figure 1, with a corresponding tilt of mobile platform, as shown in Figure 3. In motion 1, the mobile platform back side is ascending, while in motion 2, the front side is descending. The total duration of motion is set to 0.5 [s]. As results, velocity and acceleration variations of the frontal point (F) and of the center of mass (Cm) of human operator are obtained, as shown in Figure 4. International Conference on Applied Sciences Journal of Physics: Conference Series 1426 (2020) 012003 IOP Publishing doi:10.1088/1742-6596/1426/1/012003 In motion 1 (ascending of the mobile platform back side), the elevations of the characteristic points Cm and F increase, the velocities have higher values at the beginning of motion, 4459 [mm/s] - point F, and 3495 [mm/s] - point Cm, then decrease to 3345 [mm/s] - point F, and 2622 [mm/s] - point Cm; the acceleration decreases from 13641 [mm/s2] - point F and 10692 [mm/s2] - point Cm, to 4334 [mm/s2] - point F and 3397 [mm/s2] - point Cm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002563_09596518211003400-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002563_09596518211003400-Figure2-1.png",
        "caption": "Figure 2. LOS coordinates in three-dimensional space.",
        "texts": [
            " The uncertainties and their rates appeared in equation (7) have the bounded norm so that the positive unknown constants can be assumed as ri, rdi (i=1, 2, 3) which satisfy dik k\u0142 ri and _di \u0142 rdi. 2. The flight states satisfy the requirements uj j\u0142 um, fj j\u0142 fm, aj j\u0142 am, bj j\u0142 bm, and qj j\u0142 qm, wherein 0 um,fm,am,bm,qmf g p=2: 3. The items g1(q, x1), g2, and g3 are norm bounded and invertible. It can be confirmed based on assumption 2. HFV-target relative motion kinematics In the real situation, target interception happens in 3D space. According to the LOS coordinates shown in Figure 2, the HFV-target relative motion kinematics can be defined as the following nonlinear equations26 \u20acR=R _f 2 +R _u 2 cos2 f+ aTR aMR \u20acu= 2 _R _u R +2 _f _u tanf+ aTu aMu R cosf \u20acf= 2 _R _u R _u 2 sinf cosf+ aTf aMf R 8>< >: \u00f08\u00de which are developed based on the acceleration vectors of target (aTR, aTu, aTf) and the acceleration vectors of HFV (aMR, aMu, aMf) in the LOS coordinates system. Moreover, the gravity center of HFV and the coordinates origin match together and their relative motion in 3D space can be extracted using (R, u,f) parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure51.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure51.1-1.png",
        "caption": "Fig. 51.1 CAD model for the atraumatic grasper",
        "texts": [
            " After going through a thorough survey throughout India, the following areas were identified, which require innovation and are in the scope of this research. \u2022 The need for innovation in the instrument for better articulation and maneuverability. \u2022 The need to consider surgeons physical and mental comfort while performing surgery. \u2022 The need for a cost-effective solution for the Indian population. \u2022 The need for effective cleaning of the instruments. All these needs can be taken care of by innovation in laparoscopic instruments. A grasper is the first instrument that the surgeon will insert inside the abdomen during laparoscopic surgery. Figure 51.1 is the CAD model of atraumatic grasper with a double-action jaw. The end-effector consists of 12 components, including the shaft. The wristedmotion is actuated with the help of five stainless steel wires (F0.45 (mm)). The yaw part and the pitch part are responsible for producing the two wristed DoF. The range of motion of the wrist is \u00b145\u00b0 in both the perpendicular planes, and the grasper can open up to 90\u00b0. The ground part (Fig. 51.2) is completely detachable. The advantages of a detachable end-effector mechanism are: \u2022 Interchangeability: Specialized tips like grasper, scissors, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001822_tmrb.2020.3038536-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001822_tmrb.2020.3038536-Figure1-1.png",
        "caption": "Fig. 1. Interface for learning the inverse kinematics with PPO.",
        "texts": [
            " The best way is to create an user controlled instance of the task and try to solve it manually. Problems in the task definition will be recognized easily this way. To overcome these challenges, it is important to start with a simple task definition and increase the complexity progressively. The target area should be relatively large in the beginning. This section introduces the definition of the key components of the RL environment. 1) State: The arm design consists of a prismatic joint, 3 geared rolling-joints, a distal-roll, for a total of 5 DOF + gripper, as shown in Figure 1. All gear joints have the same joint limit of 50\u25e6. The q-vector is the joint angle vector. It is used to calculate the tip position and to determine the error in position and orientation. 2) Target: The target T defines the position and orientation where the tip of the arm needs to go. The target is a 4 \u00d7 4 homogeneous transformation matrix: T = ( R P 0 1 ) (2) where R is a 3 \u00d7 3 rotation matrix, P is a 3 \u00d7 1 position vector, and 0 is the vector (0, 0, 0). For training purposes, T is calculated from a random q-vector within joint limits"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.3-1.png",
        "caption": "Fig. 34.3 Structural design of the semi-span wing",
        "texts": [
            " Aerodynamic and structural limitations of the aircraft are expressed in V\u2013n diagram, also known as flight envelope. Load factor equal to 1 is the level flight condition. The curve OABE and OHGF shows the maximum velocity that an aircraft can safely operate for the corresponding load factor. Exceeding Vmax leads to structural failure as shown in Fig. 34.2. The wing is designed in 2 pieces which can be slid into the fuselage by a telescopic method. This is done to ensure easy transportation of the aircraft. As shown in Fig. 34.3, each wing has 8 ribs. An I-section passes through C/4 of the wing, an L section in the rear which not only acts as a load-bearing element but also makes it easy to add hinges and fix the ailerons. It has a leading edge of 10 mm and a trailing edge of 15 mm. A carbon fibre spar of 8 mm diameter passes through the wing which acts as the main load-bearing element. The fuselage is made of 2 sections, the front part is to house the electronics and the rear part houses the payload. The CG of payload and the CG of aircraft coincide; as a result, the plane can be flown with varying payloads or without payload"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001770_dvm49764.2020.9243883-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001770_dvm49764.2020.9243883-Figure5-1.png",
        "caption": "Fig. 5. The comparison of traditional and printed manifold",
        "texts": [
            " The hydraulic manifold that 3D-printed as a single piece (instead of previous \u2013 17 piece assembly), has less weighs by 70%, withstood the same pressure and fatigue testing (Courtesy of Penn State CIMP-3D), Fig.3. Component manufacturer Aidro Hydraulics, based in Taino (Italy), is pioneering a new approach to production, based on additive manufacturing that will revolutionize fluid power markets (Fig.4). Aidro designers made portable hydraulic power unit [4]. Components, made using 3D metal printing, were smaller and lighter than equivalent components made by conventional methods (Fig.5). At left is a manifold that had been used in a portable power unit and at right is the 3D-printed version, which weighs 74% less. Authorized licensed use limited to: Rutgers University. Downloaded on May 18,2021 at 15:37:54 UTC from IEEE Xplore. Restrictions apply. 978-1-7281-7526-3/20/$31.00 \u00a92020 IEEE Aidro created a Solutions Centre for Additive Manufacturing in Hydraulics (SCAMH) where the additive manufacturing specialists design new hydraulic solutions considering customer requirements. 3D printed parts are made using Direct Metal Laser Sintering (DMLS) technology with the EOS printer machines [5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure7-1.png",
        "caption": "Fig. 7. Pendulum axis angular measurement error.",
        "texts": [
            " The COG of the measured body is obtained by estimating the intersecting point of the pendulum axes in the RCS under different suspended postures of the body. The OdZd-axis of PCS is regarded as pendulum axis in measurement, so measurement errors of the OdZd-axis is an important cause of the COG measurement errors. In establishing the PCS, the normal direction for the face of the pendulum disk is measured and is used as the pendulum axis. Because of measurement errors of the normal direction (measured pendulum axis), the dotted line shown in Fig. 7, is not the same as theoretical pendulum axis, the continuous line shown in Fig. 7. Assuming the angle between these two axes is \u03b8. The RCS of measured body B1 is Or-XrYrZr as shown in Fig. 7. In the measurement, the changing suspended posture is the same as changing position of pendulum axis in RCS from view point of geometry analysis. Under two suspended postures, the measured pendulum axis and the theoretical pendulum axis are assumed both on plan Or-XrYr, and the origin points of PCS are represented by O1 and O2, respectively. The coordinates of O1 and O2 in RCS are assumed as (x1, y1, 0) and (x2, y2, 0), respectively. The theoretical COG (C1) and measured COG (C2) of the measured body are obtained by intersecting pendulum axes in RCS for two measurement as shown in Fig. 7. Assuming gradients of lines O1C1 and line O2C1 are k1 and k3 respectively, then gradients of lines O1C2 and line O2C2 are calculated using: { k2 = tan ( tan\u2212 1k1 \u2212 \u03b8 ) k4 = tan ( tan\u2212 1k3 \u2212 \u03b8 ) (19) Equations for the line O1C1 and the line O2C1 are written as: { k1(x \u2212 x1) \u2212 (y \u2212 y1) = 0 k3(x \u2212 x2) \u2212 (y \u2212 y2) = 0 (20) Coordinates of C1 are calculated using Eq. (20): \u23a7 \u23aa \u23aa\u23a8 \u23aa \u23aa\u23a9 xc1 = y2 \u2212 y1 + k1x1 \u2212 k3x2 k1 \u2212 k3 yc1 = k1y2 \u2212 k3y1 + k1k3(x1 \u2212 x2) k1 \u2212 k3 (21) Likewise, the equations for the lines O1C2 and the line O2C2 are written as: { k4(x \u2212 x2) \u2212 (y \u2212 y2) = 0 k2(x \u2212 x1) \u2212 (y \u2212 y1) = 0 (22) And the coordinates of C2 are obtained by Eq",
            " From the analysis mentioned above, it is seen that if the angle between two pendulum axes under two suspended postures of measured body in RCS are closer to 90\u25e6, the measurement errors of COG will be smaller when measurement errors of pendulum axis are existed. (2) Suspended posture versus measured error of pendulum axis The measured errors of COG can be calculated using Eq. (26a), where d is the square of distance between the two origin points of the PCS in the RCS under two suspended postures, and it is related to the shape and size of the measured body. The d is considered as a constant in the following discussion. Assuming that the angle between the two pendulum axis in the RCS under two suspended postures is \u03c6, as shown in Fig. 7, then the relation between k1 and k3 is: tan\u2212 1k3 + \u03c6 = tan\u2212 1k1 (30) Eq. (30) can be rewritten as: k1 = k3 + tan\u03c6 1 \u2212 k3tan\u03c6 (31) Substituting Eq. (31) into Eq. (26a) yields: \u03942 dc = ( 1 + 1 tan2\u03c6 ) \u03b82d (32) (a) To analyze the influence of angle \u03c6 on COG measurement errors, assuming the angle \u03b8 is a constant, then an error coefficient, reflecting the relation between COG measurement errors and angle \u03c6, is defined from Eq. (32): f1 = \u03942 dc/(\u03b8 2d) (33) The smaller the f1, the higher accuracy of the COG measurement",
            " The RCS of the cuboid assembly is established as illustrated in Fig. 11, and the theoretical COG coordinates calculated from CAD model are shown in Table 3. Two factures influencing COG measurement errors are studied. The first one is the angle (A) between measured pendulum axis and theoretical pendulum axis, which is used for evaluating measured error for pendulum axis. The second one is the average (B) of the angles between pendulum axes under each suspended posture in the RCS. The definition of factor A is shown as \u03b8 in Fig. 7, and the factor B is calculated as follows. Assuming that there are n different suspended postures of the measured body in the experiment, and ek(k = 1,2,\u22ef, n) is the direction vector of pendulum axis in the RCS under each suspended posture. The angle \u03b1ij between two pendulum axes is calculated using: \u03b1ij = arccos ei\u22c5ej |ei| \u20d2 \u20d2ej \u20d2 \u20d2 (42) where \u03b1ij \u2208 (0\u25e6 ,180\u25e6 ) i, j = 1,2,\u22ef,n and i \u2215= j. Hence, the value of factor B can be calculated according to the following equation: B = 2 n(n \u2212 1) \u2211n i=1 \u2211n j=1 \u03b1ij (i \u2215= j) (43) T",
            " Because the wire length l is identified through calibration [8], and according to the measurement accuracy of mass and photoelectric sensor for measuring oscillation period, the errors estimates of each parameter can be obtained: el=0, em=0.01kg, eT\u22480.00276s, eR=0.00031m . Taking a cylindrical body as studying example, the MOI around its central axis is 0.53 kg\u2219m2 estimated from CAD mode. The calculated maximum error dJ is equal to 0.005 kg\u2219m2 using Eq. (51) so that the maximum relative error of MOI of the cylindrical body, dJ/J, is 0.94%. Assuming the measured pendulum axis has an angle \u03b8 with the theoretical pendulum axis as shown in Fig. 7, the angle \u03b8 of pendulum axis will lead to errors in the estimated angles of \u03b1, \u03b2 and \u03b3 as shown in Fig. 4. Angular error \u03b8 results in the measurement errors of MOI. Assuming the maximum angular errors between measured pendulum axis and coordinate axes of the CCS are \u03b8 too. Hence the direction cosine of the measured pendulum axis in CCS is \u23a7 \u23a8 \u23a9 cos(\u03b1 + \u03b8) = cos\u03b1cos\u03b8 \u2212 sin\u03b1sin\u03b8 cos(\u03b2 + \u03b8) = cos\u03b2cos\u03b8 \u2212 sin\u03b2sin\u03b8 cos(\u03b3 + \u03b8) = cos\u03b3cos\u03b8 \u2212 sin\u03b3sin\u03b8 (52) If error of the measured axis, \u03b8, is small enough, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure15-1.png",
        "caption": "Fig. 15 Meshed box",
        "texts": [
            " Verification of\u00a0the\u00a0Transfer Path Model by\u00a0Simulation To verify the correctness of the dynamic vibration transfer path model, transient response analysis was performed using the finite element method (FEM), and the results were compared to those of the dynamic model. The vibration response of the double-layer box was analyzed on the simulation platform. The material was set as structural steel with a density of 7850\u00a0kg m\u22123, a Poisson\u2019s ratio of 0.3, and an elastic modulus of 200\u00a0GPa. The meshed box is shown in Fig.\u00a015. Based on the actual structure of the double-layer box, fixed support constraints were set at the bolt holes of the outer box, and dynamic load was applied to the input bearing seat, star wheel carrier, and output bearing seat. The boundary conditions of the box are shown in Fig.\u00a016. The box body was subjected to load and vibration response was generated at different positions. Four points were selected at the output bearing seat, input bearing seat, and outer box foot as the evaluation points of the vibration response, which were named as evaluation point I-1\u20134, II-1\u20134, and III-1\u20134, respectively (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure8-1.png",
        "caption": "Fig. 8. Free body diagram of a single leg forces and moments.",
        "texts": [
            " San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at Static forces are studied by considering the free body diagram of the leg. The first member is the input link (proximal link); the second member as the coupler link (distal link). lpi/cos\u03b4 and ldi/cos\u03b4 represent the proximal and distal link length for each leg i. The joint torque \u03c4i provided to the input link of each leg i produces output forces fx, fy and fz and moments m\u03c6x,m\u03c6y and m\u03c6z on the mobile platform. The forces (fi1,\u2212fi1) generated by the motor toque \u03c4i are the forces (fi2,\u2212fi2) acting along the coupler link (Fig. 8). These forces in turn produce the moment in the top platform.23 Hence, fi1 = fi2 = fi3. The forces exerted along the distal link ldi on each leg i can be related to the output forces, acting along the principal directions, as shown below: n\u2211 i=1 l\u2192di fi2 ldi = [ fx fy fz ]T 1 ldi \u23a1 \u23a2\u23a3 lxd1 lxd2 .... lxdn lyd1 lyd2 .... lydn lzd1 lzd2 .... lzdn \u23a4 \u23a5\u23a6 \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 f12 f22 . . fn2 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6= [ fx fy fz ]T (17) where lxdi correspond to the ith leg distal link\u2019s x coordinates. Summing all the moments applied to the top platform: n\u2211 i=1 \u2192 ei \u00d7 \u2192 fi3 ldi = [ m\u03d5x m\u03d5y m\u03d5z ]T https://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002987_s40964-021-00199-x-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002987_s40964-021-00199-x-Figure17-1.png",
        "caption": "Fig. 17 Calculated maps of a staircase effect and b abraded tops on the automatically oriented part",
        "texts": [
            " As an example, if stopping the calculation after 5 stagnating iterations, the algorithm would have been terminated in 9 iterations (instead of 19) with a worsening 1 3 of the fitness function equal to 0.02% . However, this result cannot be generalised as it depends on the specific combination of geometry and part objectives. The best chromosome at the end of the algorithm corresponds to Euler angles x = 262\u25e6 , y = 349\u25e6 , and z = 81\u25e6 . Figure\u00a016 illustrates this orientation. The part oriented as proposed by the system is manufactured using the same process parameters as detailed above. Figure\u00a017a shows the map of stair stepping defect in the proposed orientation. It is possible to notice that the calculated value of dst is less than 0.25 for all the elements of the cluster 1 in Fig.\u00a014 a. Specifically, dst is equal to 0.043 on the roof, 0.170 on the front deck, and between 0.001 and 0.233 on the rails of the boat. The comparison of these results with the ones presented in Sect.\u00a04.1 demonstrates that the orientation proposed by the algorithm determines a drastic reduction of stair-stepping in these regions. Such benefits 1 3 were actually observed in the manufactured part, as shown in Fig.\u00a012a. Specifically, it is worth underlining that for this range of dst the stair-stepping effect is no more visible. As discussed in 3.1, the other main objective of the PBO was to eliminate abraded top defect from the boat deck. Figure\u00a017b shows the map of predicted abraded tops. The value of da is equal to 0 for all the mesh elements, since no element has Nz = 1 . In other words, the whole part is expected to be free of abraded tops. This prediction was once again validated through the observation of the manufactured part. Particularly, Fig.\u00a018b shows a detail of the boat deck to be compared with Fig.\u00a012b. The results demonstrate the 1 3 adequacy of the method for the optimisation of the main objectives imposed through the clusters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001319_ines49302.2020.9147194-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001319_ines49302.2020.9147194-Figure1-1.png",
        "caption": "Fig. 1: Simplified scheme of the multi-tank system",
        "texts": [],
        "surrounding_texts": [
            "Index Terms\u2014LQR, Optimal control, Nonlinear dynamics, Multitank\nI. INTRODUCTION\nOptimal control is a theory which consists of designing a controller which fulfills a set of criteria, at a minimum or maximum cost defined by the designer. It is broad topic, where lots of methods are used, based on the exact application to which they are applied. One of these methods is the linearquadratic regulator (LQR) approach, which calculates a statefeedback gain matrix based on a linear dynamic model, using a quadratic cost function.\nThe main goal of the paper is to control the multi tank system due to its similarity to isotope separation columns hydrodynamics [1]. Several works were accomplished for the separation column, each of theme implying complex control strategies [2]\u2013[4]. The goal of present work is to find a simple but still effective control strategy. The chosen method is the LQR control, being applied with succes in different applications. It is widely used in the control theory field, ranging from aerial applications [5], [6], to distributed systems [7] and kinetics [8]\u2013[10].\nThe multitank system is commonly present in various industrial processes [11] such as irrigation systems [12], pressure control [13], applications in the chemical and food industry [14], [15] and so on. This particular system is provided by Inteco1, which is a company that provides training equipment for researchers, students, etc. in the fields of automation, robotics, instrumentation and process control. The multitank system consists of a variable speed pump, which is used to fill\n1http://www.inteco.com.pl/\nthe upper tank, three cascaded water reservoirs with constant and variable cross sections, which create nonlinearities in the system as well as controllable valves, which act as flow resistors and level sensors based on hydraulic level measurement. The liquid flow between tanks is due to gravity. Having this in mind, the authors were provided with: the system, drivers and software, ready to use real-time controllers and tutorials.\nThe paper is organized as follows. After this short introductory section, in Section 2 is presented the process model. Section 3 deals with the proposed control strategy, while the corresponding simulation results are in Section 4. The work ends with concluding remarks.\nII. SYSTEM MODEL\nAccording to [16], a model of the discussed system is easily calculated. Considering that this is a tank system that consists of three cascaded tanks, each with a separate valve and a pump that guarantees the input flow in the first tank, the first\n978-1-7281-1059-2/20/$31.00 \u00a92020 IEEE 000185\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 13:40:49 UTC from IEEE Xplore. Restrictions apply.",
            "equation we need to consider is Bernoulli\u2019s law of fluid energy conservation and derive the laminar outflow rate of an ideal fluid from a tank, presented in equation 1:\nqo = C \u221a H\nC = \u00b5S \u221a 2g (1)\nwhere qo is the outflow rate, H is the height of the fluid, \u00b5 is the orifice outflow coefficient, S is the output area of the orifice and g is the gravitational acceleration.\nLet\u2019s denote the input flow with q. Having this in mind and considering equation 1, we have:\ndV\ndt = q \u2212 C\n\u221a H (2)\nwhere V denotes the volume of the fluid in the tank. Considering our system which consists of three cascaded tanks, we obtain the following set of equations:\ndV1 dt = q \u2212 C1\n\u221a H1\ndV2 dt = C1\n\u221a H1 \u2212 C2 \u221a H2\ndV3 dt = C2\n\u221a H2 \u2212 C3 \u221a H3\n(3)\nHowever, we control the level of these tanks, so we write equation 3 in the following form:\ndV1 dH1 dH1 dt = q \u2212 C1H \u03b11 1 dV2 dH2 dH2 dt = C1H \u03b11 1 \u2212 C2H \u03b12 2 dV3 dH3 dH3 dt = C2H \u03b12 2 \u2212 C3H \u03b13 3\n(4)\nIn the case of laminar flow, the parameter \u03b1i i \u2208 {1, 2, 3} is always 0.5, according to the Bernoulli law, however in practice this can vary because of the liquid acceleration, valves and so on. These parameters, along with Ci i \u2208 {1, 2, 3} are determined in [16] and will be discussed later.\nEquation 4 can be taken even further by denoting dVi dHi with \u03b2i(Hi), i \u2208 {1, 2, 3}. According to [16], we obtain:\n\u03b21(H1) = aw\n\u03b22(H2) = cw + H2\nH2max bw \u03b22(H3) = w \u221a R2 \u2212 (R\u2212H3)2\n(5)\nwhere a = 0.25, b = 0.345, H2max = 0.35, w = 0.035 and R = 0.364.\nModifying equation 4 multiplying both sides with 1 \u03b2i(Hi) , i \u2208 {1, 2, 3}, results the final form of our nonlinear model:\nF1 = dH1\ndt =\n1\n\u03b21(H1) (q \u2212 C1H \u03b11 1 )\nF2 = dH2\ndt =\n1\n\u03b22(H2) (C1H\n\u03b11 1 \u2212 C2H \u03b12 2 )\nF3 = dH3\ndt =\n1\n\u03b23(H3) (C2H\n\u03b12 2 \u2212 C3H \u03b13 3 )\n(6)\nConsidering the above stated nonlinear model, we can obtain a linearized model, depending on our control scheme. We consider two control schemes: \u2022 Having the pump as the input and the valves considered\nconstant \u2022 The pump and the valves are considered as input 1) Control scheme 1: Having the nonlinear model stated in equation 6, and considering as input the flow of the pump, we obtain the following matrices for our state space model:\nx\u0307 = Ax+Bu\nA = \u2212 \u2202F1 \u2202H1 0 0 \u2202F2\n\u2202H1 \u2212 \u2202F2 \u2202H2 0\n0 \u2202F3 \u2202H2 \u2212 \u2202F3 \u2202H3  B = \u2202F1 \u2202q\n0 0\n (7)\nwhere F = ( F1 F2 F3 )T is the nonlinear model stated\nin equation 6, x = ( H1 H2 H3 )T and u = ( q 0 0 )T . Computing the derivatives, the equation 7 becomes:\nx\u0307 = Ax+Bu\nA =  \u2212 C1\u03b11 H 1\u2212\u03b11 1 \u03b21(H1) 0 0 C1\u03b11 H 1\u2212\u03b11 1 \u03b22(H2) \u2212 C2\u03b12 H 1\u2212\u03b12 2 \u03b22(H2) 0\n0 C2\u03b12\nH 1\u2212\u03b12 2 \u03b23(H3)\n\u2212 C3\u03b13\nH 1\u2212\u03b13 3 \u03b23(H3)  B =  1 \u03b21(H1)\n0 0  (8)\n2) Control scheme 2: Having the nonlinear model stated in equation 6, and considering as input the flow of the pump as well as the opening of the valves, we obtain the following matrices for our state space model:\nx\u0307 = Ax+Bu\nA = 0 0 0 0 0 0 0 0 0 \nB =  1 \u03b21(H1) \u2212 H \u03b11 1 \u03b21(H1) 0 0 0 H \u03b11 1 \u03b22(H2) \u2212 H \u03b12 2 \u03b22(H2) 0\n0 0 H \u03b12 2\n\u03b23(H3) \u2212 H\n\u03b13 3\n\u03b23(H3)\n (9)\nIII. PROPOSED CONTROL STRATEGY\nThe authors propose a fairly simple control strategy for this system. Considering a real case scenario with enough computational power, one can compute a linear controller based on the linear model, at each sample. The beforementioned models are presented in a form that is suitable for this case.\nA block diagram of the system can be seen on Figure 2.\n000186\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 13:40:49 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_63_0001393_s11665-020-05019-x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001393_s11665-020-05019-x-Figure1-1.png",
        "caption": "Fig. 1 (a) Experimental setup for ND at NRSF-2. Note that in this figure, the hoop component of strain is being measured, (b) the axial component of strain is being measured, (c) and (d) schematic that shows the geometry of the A383 engine block",
        "texts": [
            " Powder standards (Fe, Ge, Ca, Mo, and Ni) were measured to quantify the 2h offset, and out-of-plane position of each detector to account for the Debye cone effect induced peak shift. Statistical analysis of calibration data suggested that the highest achievable strain resolution was 50 microstrain (Ref 9). The aluminum {311} reflection was chosen for the ND experiments in the present work. A gauge volume of 5 mm 9 4 mm 9 5 mm was defined. The larger gauge volume, compared with x-ray diffraction, is important because of the potentially large grain sizes, requiring a larger diffracting volume of material. The engine block was mounted to the sample stage, as shown in Fig. 1. The residual strains in the axial and hoop directions (shown in Fig. 1c) were measured in this work. The cylinder bridge is indicated by yellow arrow in Fig. 1(d). The residual strain, e, in the cylinder bridge was determined in two directions (the axial and hoop directions) using Eq 1: e \u00bc dhkl d0;hkl d0;hkl \u00f0Eq 1\u00de where dhkl is the measured d-spacing of Al {311} planes in the gauge volume for a given strain direction. The stress-free samples were extracted directly from the cylinder bridge by removing the surrounding cast iron liner. To remove the remaining residual stress at the surface of the stress-free samples, these reference samples were machined by electrical discharge machining (EDM) into a \u2018\u2018comb\u2019\u2019 geometry specimen, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001634_icra40945.2020.9197417-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001634_icra40945.2020.9197417-Figure3-1.png",
        "caption": "Fig. 3: Ackerman-steering model of robots",
        "texts": [
            " Robot Motion Model In this paper, we aim at driving a kinematically constrained, any shape mobile robot in a planar scenario. The robot is required to move towards the target location while avoiding the obstacles and fulfilling the kinematic constraints. For holonomic robots, straight paths are employed as there are no kinematic constraints on them. However, for non-holonomic robots, various kinds of path models including circular trajectories, asymptotically heading trajectories, spiral segment trajectories are used considering the kinematic constraints of the robot. In Fig. 3, a typical representation for an Ackermann drive robot is presented for a simple car-like robot. This robot can be imagined as a rigid body that moves in a plane that has a state space denoted by C = R2XS1. A state configuration for the robot is denoted by p = (x,y,\u03b8) as shown in Fig. 3. The rigid body frame of the robot places the origin at the center of the rear axle, with its x-axis pointing along \u03b8 . Let v denote the forward speed of the robot, L denote the distance from the front to rear axles and \u03c6 denote the steering angle of the front wheels. The non-holonomic nature of the robot is related to a kinematic constraint that does not allow the robot to move laterally. This imposes a condition, dy/dx= tan\u03b8 which can be solved as: (dy/dt)cos\u03b8 \u2212 (dx/dt)sin\u03b8 = 0 x\u0307 = dx/dt = vcos\u03b8 ; y\u0307 = dy/dt = vsin\u03b8 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002009_tro.2020.3033721-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002009_tro.2020.3033721-Figure1-1.png",
        "caption": "Fig. 1. Schematic drawing of the demonstrator setup and the corresponding (right-handed) coordinate frames.",
        "texts": [
            " traversing a tool along a curved path on a freeform 3-D surface with defined kinematic constraints. The proposed solution can be used for different applications in the production of textiles, apparel, and consumer goods, in the industry of fiber reinforced plastics for the lay-up of technical textiles, and in the packaging industry. This section introduces the demonstrator setup and gives an overview of the tape application process. Moreover, the mathematical model of the robot used in this work is shortly summarized. The demonstrator setup, depicted schematically in Fig. 1, consists of the robot KUKA LBR iiwa 14 R820, equipped with the six-axis force/torque (F/T) sensor ATI Mini 40, a desktop computer and a passive, wall-mounted tape application tool with a compliant draping roll. Additionally, a 3-D object is mounted on the tool side of the F/T sensor. The robot and the F/T sensor are interfaced via two network interface cards using the ETHERCAT protocol. The controller is implemented as MATLAB/SIMULINK model, which is executed via the real-time automation software BECKHOFF TWINCAT",
            " 4) Preparation and execution: The pre-cut tape is placed in the required starting position on the feed of the application Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 17,2021 at 10:13:04 UTC from IEEE Xplore. Restrictions apply. tool and the robot performs the impedance-controlled tape application process. The demonstrator setup is described mathematically via the kinematic relations between the coordinate frames, and the dynamic robot model of the KUKA LBR iiwa 14 R820. 1) Coordinate Frames: The (right-handed) coordinate frames in Fig. 1 are denoted by calligraphic letters. The robot base frame B and the tool frame T reside stationary in the inertial (world) reference frameW . The y-axis of T is aligned with the draping roll axis. The target 3-D object is described in the object frame O. This frame coincides with the robot end-effector frame E , which is attached to the tool side of the F/T sensor. The contact frame C describes the position and orientation of the contact point of the draping roll along the path. The tape in the planar state is described in a separate coordinate system A",
            ", [17] and [20], the coordinate transformation results from a parallel transport frame, which evolves from a system of differential-algebraic equations along the path. Thus, the surface normal vector can only be specified at the initial position on the path, whereas surface-based path following control presented in this article allows to fix the surface normal vector at each position along the path. Similarly, the FRENET\u2013SERRET frame used in [15] derives from the path geometry directly and does not allow to specify a surface normal vector. As shown in Fig. 1, the demonstrator setup uses a fixed tool frame T , whereas the path is attached to the 3-D object in the object frame O, which moves with the end-effector frame E . In terms of homogeneous transformations [25], the kinematic relations are described as HTO(q) = HEOH B E (q)H W B HTW , (17) where HYX is the homogeneous representation of the coordinate frameY with respect toX , expressed inX . In this setup, E = O and thus, HEO is equal to the identity matrix. Based on (17), the output y \u2208 R6 of the system (1) is chosen as Cartesian position yp \u2208 R3 and orientation yo \u2208 R3 of the tool frame T with respect to the object frame O y = [ yp yo ] = [ hp(q) ho(q) ] = h(q) ",
            " This way, the contact point between the target 3-D object and the draping roll as well as the robot elbow position can be adjusted automatically by the nullspace controller. The desired robot motion in the nullspace is then specified by suitable nullspace objective functions [25] and the respective controller parameters. Thus, trajectory planning for the tape application process is reduced to finding a feasible robot starting pose as described in Section II-B. Note that the details of the nullspace control are beyond the scope of this article. V. IMPLEMENTATION In this section, the proposed control approach is applied to the demonstrator setup depicted in Fig. 1. The controllers are implemented as MATLAB/SIMULINK modules, which are executed via the real-time automation software BECKHOFF TWINCAT on a desktop computer. For the practical implementation, a spatial discretization of the 3-D objects must be performed and the control algorithms have to be implemented in discrete time with the sampling time Ts. In the following, the indexk of a quantity refers to the time t = kTs, k = 0, 1, 2, . . . , i.e. \u03bek = \u03be(t)|t=kTs . In this section, discretespace paths and surfaces are introduced first and based on this, the discrete-time version of the surface-based path following control is derived"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002963_16878140211028448-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002963_16878140211028448-Figure1-1.png",
        "caption": "Figure 1. (a) BHR-6 humanoid robot, (b) BHR-6 model, and (c) simplified structure: the three-linkage inverted pendulum of the take-off phase.",
        "texts": [
            " For better modeling of the vertical jump of the humanoid robot, this paper puts forward several reasonable assumptions as follows: (AS1) The whole vertical jump process is completed in one plane. (AS2) The robot is completely symmetric with respect to the sagittal plane. (AS3) In the take-off phase, there is no slipping between the robot\u2019s feet and the ground. (This is satisfied when the ground friction constraint and ZMP constraint are satisfied, which will be discussed in detail later.) (AS4) Each joint is individually driven by a motor. According to AS1-2, the BHR-6 model can be simplified to the sagittal plane model. Thus, as shown in Figure 1, the simplified structure of an open-loop kinematic chain composed of four links and three joints are obtained. Linki (i=0, 1, 2, 3) is the link of foot, shank, thigh, and the upper body of the humanoid robot, respectively. di(i=0, 1, 2) denotes the length of the foot, the horizontal distance from the ankle to the heel and the toe, respectively. l0 is the distance from the ankle to the sole of the foot, and m0 is the mass of the foot. li(i=1, 2, 3), mi(i=1, 2, 3) represents the length and the mass of Linki, respectively. The CoM of each link is located at its geometric center. As a sagittal model, m3 is half the mass of the actual robot\u2019s upper body. In the take-off phase of the vertical jump of the robot, since there is no relative movement between the foot and the ground, the simplified model of jump movement in Figure 1 can be further simplified to a threelinkage inverted pendulum fixed on the ground. As shown in Figure 1, the origin of the world coordinate system O XY is established at the robot\u2019s ankle joint, with the Y-axis vertically upward and the X-axis pointing to the front and remaining horizontal in the sagittal plane of the robot. According to the coordinate system O XY , the generalized coordinate system Qs = q1, s; q2, s; q3, s\u00f0 \u00de, as shown in Figure 1, is defined, and counterclockwise is positive. Defined by the coordinate system, q2, s is positive, q1, s and q3, s are negative, and the initial state of the robot is vertical standing state. According to the Lagrange equation, the dynamic equation of the take-off phase is expressed as equation (1): Ds Qs\u00f0 \u00de\u20acQs +Cs Qs, _Qs Qs : +Gs Qs\u00f0 \u00de=Bsus \u00f01\u00de Where Ds is the inertia matrix, Cs is the Coriolis and centrifugal force term, and Gs is the gravity term in the generalized coordinate system Qs. us = t1, s; t2, s; t3, s\u00f0 \u00de represents the torque of the ankle, knee, and hip, respectively",
            " Kt, Jt, and b denote the torque constant, the moment of inertia, and damping coefficient of the motor, respectively. Fv and Fs are the friction coefficients of the reducer. um and uj are the speeds of the motor and the joint, respectively. ic is the current and N is the gear ratio. To determine the time of the robot leaving the ground, it is necessary to calculate the acceleration of the CoM of the robot in the vertical direction. When it is equal to g, there is no longer any interaction between the robot\u2019s foot and the ground, that is, the robot enters the air. From Figure 1, the position of the CoM of the robot in the take-off phase can be calculated as PCoM , s = cx, s; cy, s . Furthermore, the acceleration of the CoM is obtained by second derivative as equation (4): \u20acPCoM , s = \u20accx, s,\u20accy, s = d dt \u2202Pcom, s \u2202Qs _Qs + \u2202Pcom, s \u2202Qs \u20acQs \u00f04\u00de In the take-off phase, the CoM of the humanoid robot is expected to be accelerated by high torque of the joints, and at the same time obtains a high initial speed for the flight phase. In this section, based on the three-linkage inverted pendulum model of the take-off phase proposed in Section 2, the motor at the joints can do work sufficiently as far as possible to achieve stable and efficient vertical jump under the conditions of robot ZMP constraints and others",
            "2m high platform. After that, a simple jump experimental platform is designed and made for preliminary experiment with a control system based on the EtherCAT bus. Then, in the experiment, the effectiveness of the motion coordination strategy proposed is verified. In this section, MATLAB and V-rep are adopted to establish the dynamic simulation environment. The controller is written in MATLAB in terms of the control flow chart in Figure 4. Considering the simplified model of jump movement model in Figure 1, the BHR6 model is built in V-rep. The dynamic parameters of simplified BHR-6 model are set as shown in Table 1. In addition, the foot parameters of BHR-6 model are shown in Table 2. Every period, MATLAB Controller sends a set of joint control parameters to V-rep for real-time simulation, and gets the joint state parameters returned. The BHR-6 model is transformed from a standing upright position to a squat initial take-off posture through the PVT interpolation. According to the control flow chart that we mentioned before, the motion trajectory of the model of every period can be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002095_012032-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002095_012032-Figure6-1.png",
        "caption": "Figure 6. Model of the laminated composite material in bending with an applicator at the middle of the span",
        "texts": [
            "1088/1757-899X/998/1/012032 A parameterized model with the finite element of the two laminates were constructed in Abaqus software to perform the simulation of mechanical behavior such as Stresses, strain, and displacements in both tension test and bending test. Here, the results attained by FEA in the bending test would compare with experimental results for the elastic domain only. It is shown that the composite layup tool for shell-type parts has facilitated to create stacking sequence and ply orientation as per requirement. It was designed a shell whose dimensions are 127 x 20 mm in bending and 165 x 20 mm in the tensile test. The scheme of loading & mesh model in both bending and tensile test was as shown in figure 6 and figure 7. In tensile test, Laminate-1 employed of maximum load of 4281 N with total elements about 672, Laminate-2 subjected to maximum load of 4030 N with 656 total elements and for bending test maximum force applied for elastic domain about 100N for both specimen composites. In this paper, the finite element model was defined with a 4-node doubly curved S4R element which was used for thin or thick shell element for both tests. Boundary conditions were applied according to loading for tensile and bending test separately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000129_978-3-319-64301-4_20-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000129_978-3-319-64301-4_20-Figure8-1.png",
        "caption": "Fig. 8 Screenshot of Skills-o-mat: a serious game for alginate mixing skill training with the help of wearable surface EMG sensor",
        "texts": [
            " Installing Using the Internet of Things for Enhanced Support of Workers in Manufacturing 161 all parts correctly requires that the worker knows that a functional item number ending on an odd digit relates to the left-hand side, whereas an even number relates to the right-hand side. In order for the adwisar system to support a work process, the process has to be formally described. An established standard for modeling processes is the Business Process Model and Notation, BPMN (Object Management Group 2011). Figure 8 shows the graphical representation of the BPMN model describing the process of installing side panel 2. The process begins at the right-hand side, at the \u201cStart\u201d node. Each square node (called tasks) contains information relevant to the worker, either a description of an activity to perform or other information about an error or warning, etc., displayed in a mobile application. After the worker has confirmed that he performed the action or read the information, the system proceeds to the next node",
            " (Reprint from Kutafina et al. 2016) the mixing technique (McDaniel et al. 2013). A serious game called Skills-o-mat developed at RWTH Aachen University (Hannig et al. 2013) is designed to provide an encouraging and entertaining way to improve the skill of alginate mixing. The training is done by simultaneously watching an instructional video and performing the mixing. One Myo armband worn by the player sends the data to a machine learning module, which returns a real-time feedback on the performance quality (Fig. 8). The application also provides summary feedback accompanied by achieved virtual \u201cpoints\u201d and \u201cmedals.\u201d Multiple research groups quickly noticed the potential of low-cost gesture recognition with the Myo armband. The level of technical knowledge on the data collection and processing is constantly increasing and enables the work on better and more advanced wearable enhanced learning platforms. In this section we take a look on some recent advances in gesture recognition with the Myo. Proprioceptive neuromuscular facilitation (PNF) is a common treatment concept of physiotherapy, focused on increasing the range of motions (Hindle et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000103_lra.2019.2952323-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000103_lra.2019.2952323-Figure3-1.png",
        "caption": "Fig. 3. Left (a): the initial posture. Right (b): the posture when establishing the hand contact with the wall.",
        "texts": [
            " We used the Choreonoid environment with the OpenHRP3 physical engine [39]. The simulation model was derived from a small-size humanoid robot HOAP-2 (height 0.5 m, weight 6 kg) [40], refurbished with seven-DoF arms. In all simulations, the left hand was used to clean a vertical wall. Using the RAA control input, a right arm swinging was generated to minimize the vertical (yaw) component of the rate of change of the system angular momentum and thus, of the yaw moment at the feet. The initial configuration is shown in Fig. 3 (a). Fifth-order splines were used to generate the desired motion (swing foot and hand) and force (hand) trajectories. The robot performed three steps straight forward in the x direction, starting with a zero CoM position. The final CoM position was set at 0.24 m in the same direction. The step length and height were set at 0.08 m and 0.02 m, respectively. The offsets were set at xvrp,T = 0.025 m and xvrp,H = \u22120.015 m (w.r.t. the ankle center)3. The duration of the step was Tstep = 1.0 s, the single and double stances took 0",
            " To maintain the initial upright posture, the desired angular acceleration and velocity in \u03c9\u0307 ref B were set at zero. The P and D feedback gains were the same as for the swing leg. This resulted in a stiff trunk. The damping gain for the joint motion was D\u03b8 = 100. The two plots of the ground reaction forces (GRFs) and the footprint plot with the ZMP and the ground projection of the DCM are shown in the left column (a)-(c) of Fig. 4. The same parameters and feedback gains as above were used for the walking control component. The left hand was commanded to contact the wall at the height of 0.4 m (see Fig. 3 (b)). The final posture was set slightly beyond the wall s.t. the hand motion always continued until contact was established. The contact itself was detected by the wristmounted six-axis force/torque sensor. Note that with this setting, the velocity just before contacting the wall was quite low. Thus, a smooth motion-to-contact transition could be ensured. The hand contact was set as frictionless. The hand motion along the unconstrained directions (x \u2212 z translation and y rotation) was determined from the feedforward plus PD feedback control laws mentioned in Section II-B1",
            " The effect of the right-arm swinging is shown in Fig. 6. It is apparent that the arm swinging contributed to the minimization of the z component (the blue line) of the system angular momentum. This resulted in the minimization of the yaw moment at the feet [27]. We also confirmed that such minimization can be achieved with active pelvis rotation, as shown in the video clip accompanying this work. In the final simulation, a disturbance force was applied in the x direction at a point around the neck (see Fig. 3 (b))4, while the robot cleaned the wall. We first examined the effect of a continuous and persistent disturbance, as shown in Fig. 5 (c). The effect of the disturbance is apparent from the increased GRFs in the x direction (the red lines in Fig. 4 (g)) when compared to the respective GRFs in Fig. 4 (d) (no disturbance). The effect is also apparent when comparing the two footprint plots in Fig. 4 (f) and (i); the DCM in Fig. 4 (i) is displaced towards the acting disturbance. Note that the ZMP trajectory did not change significantly, thus the performance was stable",
            " We also added a video clip to demonstrate the capability of the controller to deal with not-flat surfaces. We also confirmed the robustness of the controller w.r.t. a disturbance-estimation error. We have shown that robust walking along the originally preplanned footholds is possible under the presence of external perturbations and/or the hand reactions. To this end, we introduced the concept of the CoM/VRP compliance and 4The exact coordinates were (0,\u22120.05, 0.145) m in the base-link reference frame, see Fig. 3 (a). 5Note that the compliance in the trunk did not alter the behavior when the disturbance was a continuous one. 2377-3766 (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. confirmed the usefulness of trunk compliance in the case of impact-type disturbances. We have also shown that, in addition to smooth contact transitions, our controller can ensure hand motion/force control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002987_s40964-021-00199-x-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002987_s40964-021-00199-x-Figure18-1.png",
        "caption": "Fig. 18 Pictures of the part printed in the orientation suggested by the system",
        "texts": [
            " Specifically, it is worth underlining that for this range of dst the stair-stepping effect is no more visible. As discussed in 3.1, the other main objective of the PBO was to eliminate abraded top defect from the boat deck. Figure\u00a017b shows the map of predicted abraded tops. The value of da is equal to 0 for all the mesh elements, since no element has Nz = 1 . In other words, the whole part is expected to be free of abraded tops. This prediction was once again validated through the observation of the manufactured part. Particularly, Fig.\u00a018b shows a detail of the boat deck to be compared with Fig.\u00a012b. The results demonstrate the 1 3 adequacy of the method for the optimisation of the main objectives imposed through the clusters. Since for all the mesh elements Nz \u2260 1 , dc is equal to 0 everywhere, i.e., capillarity is not expected to affect the proposed solution. In fact, this defect was not observed on the manufactured part. On the other hand, the calculation of the descriptor dth points out regions highly exposed to the risk of thermal bleeding. Figure\u00a019 shows the predicted map of this defect, which can be actually observed on the manufactured part in Fig.\u00a018. Particularly, the figure points out that the calculated values of dth are between 0.782 and 0.833 in correspondence to the areas where the defect was found. As thermal bleeding was less relevant than other objectives imposed by the user for the optimisation, the intelligent system permitted these regions. As a further example of this, the part in its final orientation presented stair-stepping on the tail of the boat, which was included in the third cluster. These defects demonstrate that the final solution is a trade-off between the different requirements based on the priorities defined by the user, as discussed in Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002881_s42417-021-00329-3-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002881_s42417-021-00329-3-Figure6-1.png",
        "caption": "Fig. 6 Dynamic model and bearing form of deep groove ball bearing. a Odd pressure load. b Even pressure load",
        "texts": [
            " Besides, the asynchronous rotation of the inner and outer rings of the bearing during the operation of the system will cause the rotation and revolution of the balls. It will cause the dynamic change of the roller bearing form. At the same time, due to the existence of elastic contact deformation, the load change will cause the dynamic fluctuation of the bearing displacement. The roller load distribution of two different forms of the deep groove ball bearing during the operation is presented in Fig.\u00a06. And Fig.\u00a06a shows the odd pressure load when the radial load is mainly at the lowest roller. Meanwhile, Fig.\u00a06b shows the even pressure load when the radial load is mainly at the bottom two rollers. The basic parameters of the bearing are given in Table\u00a03. The calculation expression of dynamic radial displacement during bearing operation is: where Qmax is the maximum normal load of the roller. D is the diameter of the roller. b is the contact angle of the bearing. The bearing stiffness at odd pressure is: (6) r(t) = 4.36 \u00d7 10\u22124 Q 2\u22153 max D1\u22153 cos b where Fr is for the radial load of the bearing. In the case of even pressure, the load at the roller azimuth angle = \u00b122"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000431_ab6158-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000431_ab6158-Figure7-1.png",
        "caption": "Figure 7.Model of the system used for the numerical analysis, showing the arrangement of devices and the shape of the third vibration mode.",
        "texts": [
            " Then, wireless sensors have been positioned as a function of the nonnodal positions of all modeled modes, focusing on the one excited by the disturbance force. The aim was to satisfy equation (20), which is relatively easy for simple systems (like the one used), but which may become very difficult for complex systems. To overcome this problem, the developed device can perform a preliminary modal identification in order to check the controllability of the system; if the device is badly located, the board will communicate the information to the operator and then it will be possible to move the actuator in a better position. In figure 7 is shown the model of the system, which represents the physical arrangement of devices with respect to the beam (the system is excited at the frequency of mode III); devices locations placed on the beam are summarized in table 3. As explained before, the state recovery is triggered in each device when the error between the estimate and the local measurement exceeds a certain threshold (\u00f2max); therefore, by varying the value of this threshold, the frequency with which the data are exchanged within the network can increase or decrease"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000143_0954406219884971-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000143_0954406219884971-Figure6-1.png",
        "caption": "Figure 6. The coordinate sketch map: (a) of EMFi sensor array and (b) journal bearing.",
        "texts": [
            " The amplification factor of charge amplifier for each channel is chosen as 100mv=pC. The voltage data are acquired through the data acquisition blocks (NI 9229). The data are sampled at a rate of 10 kHz per channel. The oil film forces acting on measurement points can be obtained through dividing the acquired voltages by amplification factor and sensitivity of EMFi sensor array. The acquired signals should be further processed to obtain the pressure distribution of test bearing. The layout and coordinate system of EMFi sensor array are shown in Figure 6(a). The edges of the sensor array correspond to the boundary of journal bush, which means that the dynamic pressure at the edges was 0. The pressure distribution equation can be assumed as follows P \u00bc X4 i\u00bc1 X4 j\u00bc1 Aij sin i\u2019 1 2l L 2 \" # l j 1 \u00f02\u00de where L is the length of EMFi sensor array and \u00bcW=r is the angle which EMFi sensor array covered along the bearing pads. W denotes the width of EMFi sensor array and r is the radius of journal bearing. Aij denotes the undetermined coefficient. The oil film force acting on kth (k\u00bc 1 16) measurement point can be expressed as Fk \u00bc Z lk2 lk1 Z \u2019k2 \u2019k1 Pr d\u2019dl \u00bc X4 i\u00bc1 X4 j\u00bc1 Aij \u00f0k\u00de ij \u00f03\u00de where \u00f0k\u00deij \u00bc r i cos i\u2019k1 cos i\u2019k2 l jk2 l jk1 j l j\u00fe2k2 l j\u00fe2k1 j\u00fe 2 2 L 2 \" # is determined by the coordinate values of kth measurement point shown in Figure 6. To obtain the coordinate values of each measurement point, more details of size and shape of designed EMFi sensor array are shown in Figure 6(a) and Table 2. The oil film force Fk in Eq. (3) can be measured through EMFi sensor array. Equation (3) can be further written as equation (4), which is used to obtain the parameter Aij. With the help of Aij, the expression of pressure distribution can be built and the oil film force acting on journal at horizon and vertical direction can also be calculated through equation (5). Fx \u00bc Z L=2 L=2 Z 0 P1r sin \u2019\u00fe 1\u00f0 \u00ded\u2019dl Z L=2 L=2 Z 0 P2r sin \u2019\u00fe 2\u00f0 \u00ded\u2019dl Fy \u00bc Z L=2 L=2 Z 0 P1r cos \u2019\u00fe 1\u00f0 \u00ded\u2019dl Z L=2 L=2 Z 0 P2r cos \u2019\u00fe 2\u00f0 \u00ded\u2019 dl 8>>>>>>>>>>>>>>< >>>>>>>>>>>>>>: \u00f05\u00de where Pm(m\u00bc 1,2) denotes the pressure on mth bearing pad which is shown in Figure 6(b). 1 and 2 are the angels between vertical direction (axial y) and originals of local coordinate system on pads. The dynamic pressure and forces are tested at several different rotational speeds. During the tests, the oil was supplied at approximately 25 C and 0.2\u20130.3MPa. The displacements of journal are detected using eddy current displacement sensors settled in the test rig. The measured displacements on horizon(x) and vertical(y) directions as well as the obit are shown in Figure 7(a) and (b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003011_j.ijepes.2021.107318-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003011_j.ijepes.2021.107318-Figure3-1.png",
        "caption": "Fig. 3. Model of SISC.",
        "texts": [
            " Therefore, synchronous condenser needs to work in both under-excitation state and over-excitation state, which is different from traditional synchronous generators. In healthy condition, the air gap magnetic field is symmetrically distributed. The schematic diagrams of magnetomotive force (MMF) vectors in under-excitation state and over-excitation state are shown in Fig. 2 (a) and Fig. 2 (b) respectively. As SISC happens, a new loop with a large short-circuit current id will be generated. Synchronous condenser has two parallel branches in each phase. Taking the SISC in one branch of phase A as an example, the SISC model is shown in Fig. 3. The armature response magnetic field in normal condition is a rotating magnetic field that rotates synchronously with the rotor. Once SISC occurs, short-circuit current id will induce a pulsating MMF which distributes in space following the cosine-function M. Jiang et al. International Journal of Electrical Power and Energy Systems 133 (2021) 107318 regularity and written as: fd(\u03b1, t) = Fdcos(\u03c9t)cos(\u03b1) (1) where the Fd denotes the magnitude of pulsating MMF, \u03c9 is the electrical angular frequency, and \u03b1 is the mechanical angle of stator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000780_cdc40024.2019.9028901-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000780_cdc40024.2019.9028901-Figure2-1.png",
        "caption": "Fig. 2. Cooperative spectrum sensing procedure.",
        "texts": [
            " However, several investigations pointed out that sensing carried out locally by single devices is not accurate enough for the safe coexistence of primary and secondary users [44]\u2013[46]. Thus, it is generally agreed that one of the ways to increase the reliability of spectrum sensing is to apply cooperation between nodes. In cooperative spectrum sensing every node in a cognitive network senses the spectrum, and reports local sensing results, which are then used for acquiring a global decision characterised by the global probability of detection (see Fig. 2). In the following subsections each of these phases will be discussed in detail. The goal of spectrum sensing is for the cognitive entities (secondary users \u2013 SU) to decide on the presence of a licensed-system signal (a PU signal). If a SU detects signal r(x), the spectrum sensing decision D(x) can be treated as a double-hypothesis statistic test: D(x) = { H0 if r(x) = n(x) H1 if r(x) = s(x) + n(x), (1) where H0 is the hypothesis that the received signal is just noise n(x), i.e., the frequency band is vacant, and H1 is the hypothesis that the r(x) is the sum of noise and PU signal s(x), i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002017_cce50788.2020.9299157-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002017_cce50788.2020.9299157-Figure1-1.png",
        "caption": "Fig. 1. CAD Design of the Dual-System.",
        "texts": [
            " Finally, in section V, conclusions are presented and future work is established. The prototype was adopted from a commercially available platform due to its proven ease of construction and rigidity once built, the FT Explorer [5]. Nevertheless, the original design has been also adapted as a Vertical Takeoff and Landing (VTOL) platform. Therefore, two beams or motor pods have been added under the wing on each side of the symmetrical plane of the UAV. Each extremity of the pods will carry a motor with its respective propeller, comprising the multi rotor system. Fig. 1 shows the aircraft modeled using Autodesk\u00ae Inventor\u00ae Computer-Aided Design (CAD) software. As previously stated, its functioning comprises two systems that are operationally separated. During takeoff, only the multi rotor system is turned on and motion is controlled by the differential thrust from the lifting rotors. Yaw, roll and pitch control is similar to that of a quadrotor. For the transition phase from hovering to horizontal flight, the pusher rotor is used to give forward acceleration and gradually reach the desired cruising speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002095_012032-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002095_012032-Figure5-1.png",
        "caption": "Figure 5. Defining the layout layers (Stacking sequence) in case of two composite materials theoretically analyzed: (a) Laminate-1, (b) Laminate-2.",
        "texts": [
            " Specimen model was created in part module and simulation carried out by given input data [12] as shown in table.2 *taken from ANSYS 16.0 material library. A deformable 3D model created in dog bone shape as per ASTM D638 in tensile test. Composite Layup option makes facilitate to create required two laminates simply stacking sequence and assign each ply orientation as per design in software i.e. 00 and 450 as shown in figure (4a) & (4b) and stacking arrangement of two type laminates plies shown in figure 5. 3rd International Congress on Advances in Mechanical Sciences IOP Conf. Series: Materials Science and Engineering 998 (2020) 012032 IOP Publishing doi:10.1088/1757-899X/998/1/012032 A parameterized model with the finite element of the two laminates were constructed in Abaqus software to perform the simulation of mechanical behavior such as Stresses, strain, and displacements in both tension test and bending test. Here, the results attained by FEA in the bending test would compare with experimental results for the elastic domain only"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001707_0309324720958257-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001707_0309324720958257-Figure7-1.png",
        "caption": "Figure 7. (a) Sectional view of geometric model of YS9-8 3 19 braided wire rope (b) Schematic diagram of elliptical key point and strand expansion line.",
        "texts": [
            " Mechanical response of rope strands under tension-torsion load. Due to the special structure of the braided wire rope, the strands have complicated spatial curves, resulting in a complicated state of the braided wire rope. To facilitate the analysis, it is necessary to simplify the structure of the braided wire rope. The fourcenter circular approximation ellipse method is used to simplify the left and right braided trajectories of the braided wire rope ellipse into four arcs. The relative positions of the strands are shown in Figure 7(a). For example, the right-hand braided strand 1 with an initial phase angle of 0 is selected, and the unfolded line of the strand center line is as shown in Figure 7(b). Each strand in the braided wire rope is treated as a whole. Considering the unfolding line of the jth segment of the ith strand as an example, the axial tension along the strand is Fij, shear force is Qij, component of the two forces in the axis direction of the braided wire rope is Tij, component of the horizontal direction is Pij, torque of the strand is Mij, and braiding angle after loading is b\u2019ij that is defined as the center line of the braided wire rope strand. As shown in Figure 8, i is the ith strand of the braided wire rope, i=1, 2, 3, ",
            " Within the elastic range, the relationship between the axial tension and the strain of wire is as follows: Ti =pER4ei \u00f032\u00de Substituting the strain into the above formula can obtain the axial tension of wire, and then the stress expression of wire is as follows: si = Ti Ai = pER4ei Ai \u00f033\u00de In summary, equations (29)\u2013(31) and (33) are the mechanical response models of the braided wire rope under the tension-torsion load. Mechanical response of braided wire rope under tension\u2013torsion load The braided wire rope is composed of strands that are composed of core wire and inner and outer layer side wires. The strain of the whole rope is obtained by projecting all the above elements onto the axial direction of the rope. In this study, the strain of the braided wire rope is determined using the strain projected from the core wire strain onto the axial direction of the rope. It can be seen from Figure 7 that there exists the following relationship between the strain of core wire and the axial strain of wire rope: es = e0 cosb 0 ij \u00f034\u00de , where es is the strain of the whole rope and b 0 ij is the braiding angle of the wire rope after being loaded. According to Figure 7(b), there exist two different slopes of each strand, that is, two different braiding angles. Therefore, the axial strain along the rope under the tension\u2013torsion load of the strand core within a pitch is equal to the sum of the strains of each segment, that is: es =3e0 cosb 0 i1 +2e0 cosb 0 i2 \u00f035\u00de , where subscripts 1 and 2 represent two weaving angles. It can be seen from equation (6) that the external torque of the whole braided wire rope can be calculated using the following equation: Mt = X8 i=1 Mi \u00f036\u00de Assuming that the tension of strands is equal along the axial direction of the rope, the stress of the rope can be replaced by the strand stress"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure4-1.png",
        "caption": "Fig. 4. Bearing coordinate system.",
        "texts": [
            " The coefficient nK is determined by the material and geometric structure of the bearing (raceway curvature, raceway diameter, steel ball diameter, etc.). It is assumed that the rings are rigid, that is, only contact deformation only occurs at the contact part with the steel ball. The rings\u2019 shapes are kept unchanged in the macroscopic view. So the equilibrium equations of the external load xF , yF , zF , yM , zM and the ring displacement x\u03b4 , y\u03b4 , z\u03b4 , y\u03b8 , z\u03b8 can be established. The coordinate system, the direction of external load and the direction of displacement are shown in Fig. 4. Assuming that the total number of steel ball is Z , they are evenly distributed in the 360\u00b0 range of the raceway. The distribution and numbering of the steel balls are shown in Fig. 5. i\u03c6 is the azimuth position of the i-th steel ball. When DGBB is under load, the steel ball has only one contact point each with the inner and outer rings; the geometric relationship is shown in Fig. 6. When the steel ball is in contact with the inner and outer rings at the same time without deformation, the distance of curvature center of the inner and outer raceway 0A can be expressed as, 0 i e wA r r D= + \u2212 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure10-1.png",
        "caption": "Fig. 10. Complete mechatronic assembly using both optimized (front plane, dark green) and non-optimized structures (rear plane, green).",
        "texts": [
            " The conductive traces are printed simultaneously in contact with the non-conductive housing, as depicted in Figure 9 along with the physical, 3D printed device. The assembly process consists of inserting the motor and the battery. The selected Tamyia RC300-FT-14270 DC motor represents a typical off-the-shelf component. It exhibits the application convenient low-current characteristics [23] of 25 mA at 1.5 V, as well as solid terminals which can be contacted against the conductive traces. The battery is a standard 9V battery from Q-Batteries [20] and is inserted into the battery holder, as can be seen in Figure 10. The latter is printed using Black Magic 3D Conductive Graphene filament [29], which comes with a bulk resistivity of 0.6 \u2126cm. The non-conductive part is printed with standard PLA from Sakata. We have assembled the device with both the optimized and non-optimized structures, and measured their performance. As a performance indicator, the power dissipated in the traces has been evaluated by multiplying the measured voltage drop over the traces with the measured current delivered to the DC motor. The results for optimized and non-optimzed nonconductive beds are 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000052_s1068366619050210-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000052_s1068366619050210-Figure1-1.png",
        "caption": "Fig. 1. Ball bearing cage: (a) is a general view of the bearing, and (b) are the main dimensions of the bearing.",
        "texts": [
            " Theoretical [3] and experimental [4, 5] studies of the friction moments in rolling bearings do not consider such structures. At the same time, full complement bearings are widely used in control systems of rocket engines [6] and small-sized bearings of machines and devices [7]. Objective\u2014Development and experimental verification of the method of calculating the friction moment of full-friction rolling bearings. FORMULATION OF THE PROBLEM A general view of the complete bearing assembly is shown in Fig. 1a. Such bearings are usually installed in pairs in an opposite way. Bearings have the same nominal dimensions. However, taking into account permissible deviations, the contact angles of the balls and the displacement of the rings during tightening can be determined by calculating each bearing. The main dimensions of the bearing (Fig. 1b) we accept without taking into account the deformations from the landing interference and thermal expansion under operating conditions. Here: and are the bearing dimensions, and are the diameters at the bottom of the groove of the inner and outer rings, and are the diameters of the sides of the base ends of the rings, is the diameter of the lock on the bottom end of the outer ring, and is the ball diameter. d D id oD bd bD lD wD 425 When the outer rings are tightened, all the balls are loaded evenly and the contact angles with the inner and outer rings will be the same"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001125_j.matpr.2020.05.084-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001125_j.matpr.2020.05.084-Figure1-1.png",
        "caption": "Fig. 1. 3D CAD models (a) Circular (b) Horizontal ell",
        "texts": [
            " The two critical geometrical parameters namely the included angle (h) and loading distance (d) were already studied for their impact on the auxetic behaviours and the stress distribution patterns and found to play critical roles in overall performance of the Sshaped structure [21]. However, the role of the ratio of the length of the inclined strut to the radius (rather the ratio of the length to twice the radius, L/2R, which will take care of the radial transformations on either side of the S- unit cell) of the S-shaped unit cell have not been explored so far and therefore form the part of the research reported in this paper. edings, Fig. 1 depicts the CADmodels of all three variations of the chiral S-shaped structure. All the models are designed in the SolidWorks software by joining the two curved parts with the inclined link of length (L). The circular S-shaped model in Fig. 1(a) is the parent design, however, the other two forms namely horizontal (Fig. 1 (b)) and vertical (Fig. 1(c)) ellipses, are derived by varying the L/2R ratios keeping other geometrical parameters constant as mentioned in the Table 1. The change in the L/2R ratio caused the change in the shape of S-shaped form, from circular to elliptical in both cases (Horizontal and vertical ellipse). The values of the geometrical parameters of all three models are listed in Table 1. The mechanical behaviour of all the three structural forms (circular, horizontal and elliptical S-shape) were simulated for compressive loading using the ANSYS Workbench 2019"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002546_s12541-021-00485-2-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002546_s12541-021-00485-2-Figure3-1.png",
        "caption": "Fig. 3 Dimensions of bolted specimen",
        "texts": [
            "46 \ufffd Pflat \u239e \u239f\u239f\u239f\u239f\u23a0 1\u22153 (30) KT = \u222b a\ufffd l a\ufffd c ktn(a \ufffd)da\ufffd= 4DG\u2217 (2\u2212D)\u22152a \ufffdD\u22152 l 0.5 \u222b a\ufffd l a\ufffd c a \ufffd \u2212 (1 + D)\u22152 B3da \ufffd (31)B3 = \u239b\u239c\u239c\u239c\u239c\u239d 1 \u2212 Qm flat 0.26 cot \ufffd 0.27 c - 1 \ufffd a\ufffd a\ufffd c \ufffd0.46 \ufffd Pflat \u239e\u239f\u239f\u239f\u239f\u23a0 1\u22153 (32) CT = TKT= 3.2448DG\u2217 (2\u2212D)\u22152 ( (2\u2212D)Aa D\u22c5 (2\u2212D)\u22152 Ar\u2217 )D\u22152 5 ( 1 2 )3\u22152 1.5 \u22c5 \u222b a\ufffd l a\ufffd c a\ufffd \u2212(2 + D)\u22152 \u22c5 B1da \ufffd \u22c5 \u222b a\ufffd l a\ufffd c a\ufffd \u2212(1 + D)\u22152 B3da \ufffd \u222b a\ufffd l a\ufffd c a \ufffd(1\u2212D)\u22152 ( 0.26 cot ( 0.27 ( c \u22121 ( a\ufffd c a\ufffd )D\u22121 )0.46 ))2 B2da \ufffd 1 3 To certify presented model, the bolted specimens are designed as shown in Fig.\u00a03. The bolted assembly is composed of two specimens which are connected by two M16 bolts. The material of bolted specimen is nodular cast iron (Material type is QT235 in China), and its material properties are listed in Table\u00a01. The roughness parameter of the machining contact surfaces is Ra = 1.6 m. For the purpose of obtaining the fractal parameters of the contact rough surface, a 1\u00a0mm*1\u00a0mm*1\u00a0mm test block is manufactured to measure the surface morphology, in which the material, roughness, and machining method are identical to those of the specimens in the experimental set-up"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure16-1.png",
        "caption": "Fig. 16. Schematic diagram of dimension parameters",
        "texts": [
            " It can be seen that the machine structure with 22 rotor pole and 24 slot pole has the best performance. Authorized licensed use limited to: Carleton University. Downloaded on June 05,2021 at 19:52:57 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (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. In this study, the dimension parameters shown in Fig. 16 are optimized using the multi-objective DE coupled with FEM. It can find global optimal solutions with a faster and smoother convergence, as compared to the algorithms of the adaptive Simulated Annealing and the Nelder-Mead method. Besides, compared to the Linear Programming (LP), which requires prohibitive amounts of computation time in processing units, the DE exhibits better efficiency to find global optimal solutions for complex systems with multiple variables [34]. The objective of the optimization is to pursue relatively higher torque and lower torque ripple"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002721_tec.2021.3074818-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002721_tec.2021.3074818-Figure3-1.png",
        "caption": "Fig. 3 Superposition based on a single rotor tooth of a FRPM machine with stator-PM. (a) Rotor with Ns teeth. (b) Rotor with 1st tooth only. (c) Rotor with ith tooth only.",
        "texts": [
            " 2(b)] is: \ud835\udf111(\ud835\udf03, \ud835\udefc) = \ud835\udc531(\ud835\udf03) \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (3) Then, given the space angle between the first tooth and the \ud835\udc56th tooth being 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1), the characteristic of the machine having the rotor with \ud835\udc56th tooth only, i.e., Fig. 2(c), is written as \ud835\udf11\ud835\udc56(\ud835\udf03, \ud835\udefc) = \ud835\udc53\ud835\udc56 [\ud835\udf03 + 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1)] \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (4) Therefore, the characteristic of the machine having \ud835\udc41\ud835\udc60 stator teeth, i.e. Fig. 2(a), can be expressed by the superposition of those having a stator with single tooth only: Similarly, the mechanism of superposition based on a single rotor tooth can be described as follows. The open-circuit characteristic of the machine with \ud835\udc41\ud835\udc5f rotor teeth, shown in Fig. 3(a), can be modeled by the superposition of those having a rotor with single tooth only. The position of each single rotor tooth, i.e., Figs. 3(b)-3(c), remains the same as Figs. 3(a). Besides, the stator topology also remains the same. The characteristic of the machine with 1st rotor tooth only, shown in Fig. 3(b), is: \ud835\udf111(\ud835\udf03, \ud835\udefc) = \ud835\udc53(\ud835\udf03) \u2219 \ud835\udc541(\ud835\udf03, \ud835\udefc) (6) Similarly, since the space angle between the first tooth and the \ud835\udc56th tooth is 2\ud835\udf0b \ud835\udc41\ud835\udc5f (\ud835\udc56 \u2212 1), the characteristic of the machine having the rotor with \ud835\udc56th tooth only, i.e., Fig. 3(c), is written as \ud835\udf11\ud835\udc56(\ud835\udf03, \ud835\udefc) = \ud835\udc53(\ud835\udf03) \u2219 \ud835\udc54\ud835\udc56 [\ud835\udf03, \ud835\udefc + 2\ud835\udf0b \ud835\udc41\ud835\udc5f (\ud835\udc56 \u2212 1)] (7) Therefore, the open-circuit characteristic of the machine having \ud835\udc41\ud835\udc5f rotor teeth, i.e. Fig. 3(a), can be expressed by the superposition of those having a rotor with single tooth only: Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 19,2021 at 05:25:29 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (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. In this sub-section, the air-gap flux density is modeled by the superposition of a single rotor tooth. Firstly, according to [25], the open-circuit flux density of an original FRPM machine with \ud835\udc41\ud835\udc60 stator slots and \ud835\udc41\ud835\udc5f rotor teeth can be expressed as: \ud835\udc35(\ud835\udf03, \ud835\udefc) = \u2211 \u2211 \ud835\udc35\ud835\udc5b1\ud835\udc5a1 sin[(\ud835\udc5b1\ud835\udc41\ud835\udc60 \u00b1\ud835\udc5a1\ud835\udc41\ud835\udc5f) \u221e \ud835\udc5a1=0 \u221e \ud835\udc5b1=1 \ud835\udf03 \u00b1 \ud835\udc5a1\ud835\udc41\ud835\udc5f\ud835\udefc] (9) where, \ud835\udc5b1 is the harmonic order of PM-excited MMF, \ud835\udc5a1 is the harmonic order of permeance. If use the superposition of a single rotor tooth to model the flux density, it can be derived in the following. As can be seen from Fig. 3, the space angle between every two teeth is 2\ud835\udf0b \ud835\udc41\ud835\udc5f . Then, according to (7) and (9), the flux density of the FRPM machine with \ud835\udc56th rotor tooth only [Fig. 3(c)] is: \ud835\udc35\ud835\udc60\ud835\udc50\ud835\udc56(\ud835\udf03, \ud835\udefc) = \u2211 \u2211 \ud835\udc35\ud835\udc5b2\ud835\udc5a2 sin{(\ud835\udc5b2\ud835\udc41\ud835\udc60 \u00b1\ud835\udc5a2)\ud835\udf03 \u00b1\ud835\udc5a2[\ud835\udefc + \u221e \ud835\udc5a2=0 \u221e \ud835\udc5b2=1 2\ud835\udf0b \ud835\udc41\ud835\udc5f (\ud835\udc56 \u2212 1)]} (10) where, \ud835\udc5b2 is the harmonic order of PM-excited MMF, \ud835\udc5a2 is the harmonic order of permeance. Thus, according to (8), the accumulated flux density considering \ud835\udc41\ud835\udc5f rotor teeth is: \ud835\udc35(\ud835\udf03, \ud835\udefc) = \u2211 \u2211 \u2211 \ud835\udc35\ud835\udc5b2\ud835\udc5a2 sin { \u221e \ud835\udc5a2=0 \u221e \ud835\udc5b2=1 \ud835\udc41\ud835\udc5f \ud835\udc56=1 (\ud835\udc5b2\ud835\udc41\ud835\udc60 \u00b1\ud835\udc5a2)\ud835\udf03 \u00b1 \ud835\udc5a2[\ud835\udefc + 2\ud835\udf0b \ud835\udc41\ud835\udc5f (\ud835\udc56 \u2212 1)]} (11) Letting \ud835\udc4e = 2\ud835\udf0b\ud835\udc5a2 \ud835\udc41\ud835\udc5f , \ud835\udc4f = (\ud835\udc5b2\ud835\udc41\ud835\udc60 \u00b1\ud835\udc5a2)\ud835\udf03 \u00b1 \ud835\udc5a2\ud835\udefc \u2213 2\ud835\udf0b\ud835\udc5a2 \ud835\udc41\ud835\udc5f , then: \ud835\udc35(\ud835\udf03, \ud835\udefc) = \u2211 \u2211 \ud835\udc35\ud835\udc5b2\ud835\udc5a2 (cos\ud835\udc4f \u00d7\u2211sin\ud835\udc4e\ud835\udc56 \ud835\udc41\ud835\udc5f \ud835\udc56=1 + sin\ud835\udc4f \u00d7 \u221e \ud835\udc5a2=0 \u221e \ud835\udc5b2=1 \u2211cos\ud835\udc4e\ud835\udc56 \ud835\udc41\ud835\udc5f \ud835\udc56=1 ) (12) Since { \u2211sin\ud835\udc4e\ud835\udc56 = sin \ud835\udc41\ud835\udc5f\ud835\udc4e 2 sin (\ud835\udc41\ud835\udc5f + 1)\ud835\udc4e 2 sin \ud835\udc4e 2 \ud835\udc41\ud835\udc5f \ud835\udc56=1 \u2211cos\ud835\udc4e\ud835\udc56 = sin \ud835\udc41\ud835\udc5f\ud835\udc4e 2 cos (\ud835\udc41\ud835\udc5f + 1)\ud835\udc4e 2 sin \ud835\udc4e 2 \ud835\udc41\ud835\udc5f \ud835\udc56=1 (13) Equation (12) can be simplified as: \ud835\udc35(\ud835\udf03, \ud835\udefc) = sin\ud835\udc5a2\ud835\udf0b sin \ud835\udc5a2\ud835\udf0b \ud835\udc41\ud835\udc5f \u2219 \u2211 \u2211 \ud835\udc35\ud835\udc5b2\ud835\udc5a2 \u221e \ud835\udc5a2=0 \u221e \ud835\udc5b2=1 sin [(\ud835\udc5b2\ud835\udc41\ud835\udc60 \u00b1\ud835\udc5a2)\ud835\udf03 \u00b1 \ud835\udc5a2\ud835\udefc \u2213 2\ud835\udc5a2\ud835\udf0b \ud835\udc41\ud835\udc5f + (\ud835\udc41\ud835\udc5f + 1)\ud835\udc5a2\ud835\udf0b \ud835\udc41\ud835\udc5f ] (14) As can be seen, (14) is meaningful only when the denominator sin \ud835\udc5a2\ud835\udf0b \ud835\udc41\ud835\udc5f \u2261 0 since sin\ud835\udc5a2\ud835\udf0b is always 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure5-1.png",
        "caption": "Figure 5. Mesh of the door for finite element analysis.",
        "texts": [
            " This system has a solution that is nonzero if j x2 M\u00bd \u00fe K\u00bd j \u00bc 0 (4) This relation is the characteristic equation associated to equation (1) and offers the values of the n eigenfrequencies of the system. To determine the eigenmodes of vibration it is necessary to solve the following linear system K\u00bd uf gi \u00bc x2 i M\u00bd uf gi (5) The first 10 eigenvalues are presented in the following figures and in the Appendix. The Hyperworks/ Hypermesh software suite was used to discretize the CAD model. For the calculation performed, the FEM model was composed of a number of 17,216 finite elements, QUAD4, plate type with the following properties: Young\u2019s modulus E\u00bc 1420 MPa and Poisson\u2019s ratio \u00bc 0.259 (see Figure 5). With the help of the modal analysis, the eigenfrequencies and eigenmodes of vibration for the door made of composite materials were determined. The first six eigenmodes are represented in Figures 6 to 8. In the Appendix are presented the eigenmodes 7\u201310 (Figures 19 and 20). Improving the structure made of composite materials Experimental research has shown that the deformations of doors made of composites are significantly higher than those of metal doors.20 By simple observation it is found that the metal door is significantly stiffer than the composite door"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure1-1.png",
        "caption": "Fig. 1. 3D view of the proposed concept",
        "texts": [
            "933) and lower ripple harmonics. It also supports both single-layer and double-layer concentrated winding configura- 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 1083 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 19,2021 at 14:42:00 UTC from IEEE Xplore. Restrictions apply. tion. A double layer winding is selected due to its reduced endturn length. The windings are wound around thermal plastic support structures in place of conventional laminated teeth as presented in Fig 1. The rotor has magnets in a Halbach configuration consisting of three segments per pole to circumvent the need for rotor core [1]. A non-magnetic rotor support is used to hold the magnet as shown in Fig 2. The coreless rotor minimizes the magnet loss and has zero rotor core loss. This helps to keep simple thermal management for the rotor. The considered machine is optimized to maximize the torque density and minimize the total loss while maintaining the parameters within the constraints given in Table I",
            " The detailed electromagnetic performances are not in the current scope of this paper. Since the winding supports/teeth of a slotless motor are non-magnetic, a thermal plastic material (Coolpoly D5506 [7]) with integrated cooling channel is used as winding support in the stator. The thermal plastic is an electrical insulator and thermally conducting. It is also easily moldable to a desired shape. This helps to introduce cooling channels through the winding support. The coolant tubes and their placement with respect to the heat sources are shown in Fig 1. Circulating the coolant in close-proximity of the heat source substantially improves the thermal performance of a slotless machine. Channels are positioned in a way so that it does not decrease the fill factor of the winding. Additionally, another structure made of aluminum is designed to provide the second path for the coolant flow just beneath the stator back-iron/yoke. This second path helps to extract the heat from the motor lamination. In this proposed concept, two parallel paths are provided for heat extraction out of the machine as presented in Fig 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000374_icoiact46704.2019.8938512-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000374_icoiact46704.2019.8938512-Figure1-1.png",
        "caption": "Fig. 1 Ball Bearing Parameter",
        "texts": [],
        "surrounding_texts": [
            "Wavelet transformation is a mathematical transformation used to analyze moving signals. Wavelet-based transformation method is one way to be able to be used to analyze non-stationary signals. This method is also used to detect certain events and can be used for data compression [12]. In addition, wavelet transformations can also be used to analyze non-stationary signals (i.e. signals whose frequency content varies with time), because they are related to their ability that signal characteristics can be studied locally and in detail, according to the scale they have. DWT can convert source signals into two signal classifications, namely high frequency signals with high time resolution and frequency of lace with high time resolution. This signal works on two DWT filters, namely highpass filters and lowpass filters so that the frequency of these signals can be analyzed. After passing through the filtration process, proceed with the sub-sampling operation by taking each half of the filter output. This process is called the decomposition process [13]. The decomposition process can be carried out continuously by paying attention to the level used, to produce the appropriate decomposition level These decompositions are limited by number; limited number of levels is difined by [14]: (7) , the sampling frequency and fundamental frequency respectively. Indicating that purpose diffrent number of levels for adequate analysis, it adds 2 to . Therefore for a frequency sampling of 10 kHz, the number of decomposition, advisable is of: 2 9 (8) There are several filtering (denoising) methods, but our technical depend on the DWT. We have put a zoom on details which are rich in harmonics and acceptable information about the fault. III. EXPERIMENTAL SETUP"
        ]
    },
    {
        "image_filename": "designv11_63_0000143_0954406219884971-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000143_0954406219884971-Figure3-1.png",
        "caption": "Figure 3. The structure and installation of designed EMFi sensor array: (a) the sketch of EMFi sensor array, (b) the structure of one measurement point, and (c) the sketch of installing EMFi sensor array on bearing bush.",
        "texts": [
            " The rotation speed varies from 300 r/min to 4260 r/min with the step of 360 r/min, and the first critical speed of the rotor system is 10,488 r/min. The shaft can be regarded as rigid rotor because the maximum rotation speed far below the first critical speed. The gear ratio of the gear box was 1/4. Two journal bearings were lubricated using same oil supply system. Turbine oil 46# is used as the lubricant in this test. Its kinematic viscosity is 44.6 cSt at 40 C and density is 0:86 103 kg=m3 at 15 C. The sketch of designed EMFi sensor array can be seen in Figure 3. The sensor array is designed for one pad of journal bearing and 16 (4 4) measurement points are included in the sensor array. Since the tested journal bearing is of two bearing pads and two EMFi sensor arrays are used to measure the pressure, there are totally 32 measurement points in this test. The charges conducted from each measurement points are concerned with resultant force acting on whole measurement area. That means EMFi sensor can only measure the average pressure or resultant force on each measurement point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002176_0142331220981430-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002176_0142331220981430-Figure3-1.png",
        "caption": "Figure 3. The relationship among angles gT , fi and l\u0302i.",
        "texts": [
            " ji is the bearing angle between the direction of the i-th attacker\u2019s velocity and the i-th LOS, and the term fi is the bearing angle between the direction of the target\u2019s velocity and the i-th LOS. Multi-attackers with uniform velocity are denoted as nodes V= 1, ::,N . For simplicity, a geometry is described for the i-th attacker and its neighbors in Figure 2. uii+ 1 is the angle from the i-th LOS to the i+1-th LOS, and rii+ 1 is the relative distance between the i+1-th attacker and the i-th attacker. VT is a positive constant that describes the initial speed of the target. The connection between the angles gT , fi and l\u0302i is formed in Figure 3. It is noteworthy that the attacker and the target have their accelerations perpendicular to the directions of their speeds, respectively, which means their speeds are pre-determined and the directions of the speeds change constantly. In this paper, the speed of the attacker is always lower than speed of the target. Dynamic equations of the i-th attacker are given by _Ri =Vri,Vri =VT cosfi Vi cos ji _li = Vli Ri ,Vli =VT sinfi Vi sin ji gi = ji +li, gT =fi + l\u0302i _gi = AMi Vi , _gT = AT VT uij =lj li \u00f01\u00de with l\u0302i = li p if li \u00f8 p li +p if li\\p where the subscripts i and T denote the i-th attacker and the target",
            " When multiple attackers satisfy Assumption 2, for example, i-th attacker can observe the relative distance Ri, the target velocity VT and LOS angle li, then i-th attacker\u2019s neighbour, j-th attacker (i.e. j 2 Ni), can also obtain these information such as Rj = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 i + r2 ij 2Ririj cos (li ai) q and LOS angle lj =li + arcsin ( rij sin (li ai) Rj ). As is shown in Figure 3, the j-th attacker can obtain the distance between itself and i-th attacker rij and angle aj (notice that ai =aj) through the geometric information of the communication graph. In the same pattern, the k-th attacker, which is within the detection radius of the j-th attacker (i.e. k 2 Nj), can also obtain its own relative position and speed with respect to the target. According to Assumption 1, relative position and speed with respect to the target for all the attackers are obtained. This section designs three multiple low-speed attackers intercepting a single target with high maneuverability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001674_icuas48674.2020.9213835-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001674_icuas48674.2020.9213835-Figure5-1.png",
        "caption": "Fig. 5: Principal axes of a polygon",
        "texts": [
            " According to (13), to minimize the cost function, our control law will drive all UAVs to the centroids of the Voronoi cells. After reaching the centroid, each UAV has its own Voronoi cell. For example, 15 UAVs is divided into 15 cells by the Voronoi partition algorithm, as shown in Figure 4, where each UAV has its corresponding cell. It is observed that each cell is of a polygon form. In order to cover the cell quickly, we intend to find the principal axes of each cell or polygon so that each camera can find its heading direction, as shown in Figure 5. Based on the major principal axis (X-axis), we need to find the largest rectangle (default FOV without the configuration of PTZ) of a given polygon. For example, Figure 6 shows the largest orthogonal rectangle in a given polygon. These results are used as an initial setting of 1105 Authorized licensed use limited to: Newcastle University. Downloaded on October 18,2020 at 04:30:19 UTC from IEEE Xplore. Restrictions apply. FOV. Thereafter, the distributed coverage control as shown in [21] is used to cover the corresponding cell as shown in the UAV\u2019s gradient law (6)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure15-1.png",
        "caption": "Fig. 15. Magnetic base of anti-bird thorn",
        "texts": [],
        "surrounding_texts": [
            "Take advantage of birds' consciousness of territory to attract birds to build nests by installing man-made bird cages, which can prevent birds from nesting in the transmission line hanging point with other anti-bird devices. However, the existing man-made bird cage is simplely made by wire mesh, as shown in Fig.14, which is not convenient to carry to the tower and is not conducive to large-scale installation. The new man-made bird cage can be folded into a flat surface by hinges, enhances portability and provides a great benefit to the processing of the devices. V. INVERTED UMBRELLA-SHAPED RETRACTABLE ANTI- BIRD THORN"
        ]
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure10-1.png",
        "caption": "Fig. 10. (a) Wire-Ring contact status; (b) MATRIX27 mechanism; (c) COMBIN39 mechanism.",
        "texts": [
            " Fp = \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305( k2 2\u22c5cos2(\u03b10) + k2 1\u22c5sin2(\u03b10) )\u221a \u22c5\u03b4p = kp\u22c5\u03b4p (31) kcbn = k3\u22c5kp k3+kp = k3\u22c5 \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305( k2 2\u22c5 cos2(\u03b10) + k2 1\u22c5 sin2(\u03b10 ))\u221a k3 + \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305( k2 2\u22c5 cos2(\u03b10) + k2 1\u22c5 sin2(\u03b10 ))\u221a (32) In order to introduce the MATRIX27 (defined by Eq. 30) or the COMBIN39 (defined by Eq. 32) elements in the FE model, they must be joined to the bearing rings. For such purpose, the procedure proposed by Daidie\u0301 for conventional slewing ball bearings [15] was adapted to the case under study, as explained below. Fig. 10a shows the wire-ring contact status when solid wires are used in the simulations, being coloured in orange the regions in contact. Fig. 10b and 10c show, coloured in grey, the surfaces of the rings to which the MATRIX27 and COMBIN39 are linked, respectively. These surfaces are rigidized though shell elements, so no local deformations are allowed, since contact flexibility is simulated by means of the MATRIX27 or the COMBIN39. The rigid areas of each ring are linked to a reference node, which location coincides with the reference points in Fig. 3b. The link between these nodes and the surfaces is done by means of rigid beams. Then, the reference nodes of the rigid shell structures of both rings are linked through the corresponding MATRIX27 or COMBIN39 element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002629_012013-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002629_012013-Figure4-1.png",
        "caption": "Figure 4. End cover fan to motor housing",
        "texts": [],
        "surrounding_texts": [
            "Modal Analysis is mostly dispensed to seek out natural frequency of the system and completely different mode shapes at different Eigen values [12,14]. In this work, totally different modes and Eigen values square measure calculates and premeditated between mode shapes versus natural frequencies [15,20, Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 21]. This shows the stiffness and rigidity of the material toward totally different modes. Using Modal analysis it can be captured the first 10 natural frequencies and related mode shape of each frequency for the 3 cases. Figure 7 shows the modal analysis results for case A. Each figure represents the frequency of each mode and displacement related to the frequency. Due to rotational velocity the model is exhibiting rotational vibration. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Table 1 shows the value of modal analysis for case A, B and C. According to this table Case C is having more natural frequency than other 2 cases. Figure 8 shows the relation between modes and damped frequency. Based on modal analysis results Case C (Hybrid Hemp - Flax fibre) has additional stiffness (due to high natural frequency) while compared to other two cases. Conclusion In this analysis, a study for three totally different composite materials particularly Hemp fibre, Flax fibre and Hybrid of Hemp and Flax fibre for Finite Element Analysis of electrical motor casing has been carried out. The motor casing should be rigid and better in stiffness to resist operative vibrations. 3D modelling of motor casing is built using Parametric Creo one among the fine 3D modelling software. These 3 materials are applied to motor casing as a material properties of the model and three types of analysis namely, Modal Analysis, Rotor Dynamics and Structural Transient Analysis carried out to check the stiffness, reliability towards operative frequency and durability of motor casing. As determined from modal analysis Case C (Hybrid Hemp Flax fibre) has better stiffness compared to two different fibres as shown in graph fig 7 (Modes vs Damped Frequency). Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Campbell diagram was premeditated between frequencies vs speed of the motor for predicting critical speed. The operative frequency and operative speed do not match with the critical speed and frequency. Model analysis was performed for finding the vibrational stability of the composite materials and observation Case C shows low deflection and equivalent stresses compared to different fibres. Force vs deflection curve and stress vs strain curve are premeditated to indicate the stiffness of the materials and energy vs deflection curve premeditate show of the strain energy observed, when deformation happens. As determined from all above studies Case C has higher stiffness and strength compare from different two fibres. Reference [1] Bajuri F, Norkhairunnisa Mazlan, Mohamad Ridzwan Ishak and Junichiro Imatomi, 2016, \u201cFlexural and compressive properties of hybrid kenaf/silica nanoparticles in epoxy composite\u201d, Procedia Chemistry, Vol.19, pp:955-960. [2] Korniejenko. K, Frczek. E, Pytlak. E and Adamski. M, 2016, \u201cMechanical properties of geo polymer composites reinforced with natural fibres\u201d, Procedia Engineering, Vol.151, pp:388393. [3] Lut Pil, Farida Bensadoun, Julie Pariset and Ignaas Verpoest, 2016, \u201cWhy are designers fascinated by flax and hemp fibre composites?\u201d, Composites: Part A, Vol.83, pp:193-205. [4] Shuhimi F, Mohd Fadzli Bin Abdollah, Kalam M.A and Hilmi Amiruddin, 2016, \u201cTribological characteristics comparison for oil palm fibre/epoxy and kenaf fibre/epoxy composites under dry sliding conditions\u201d, Tribology International 101. [5] Raman Bharath, Vijaya Ramnath B and Manoaran N, 2015, \u201cKenaf fibre reinforced composites: A review\u201d, ARPN Journal of Engineering and Applied Sciences, Vol.10, No.13, pp:5483-5485. [6] Zamria M, Hazizan Md Akil and Zainal Ariffin Mohd Ishak, , 2016 \u201cPultruded kenaf fibre reinforced composites: Effect of different kenaf fibre yarn tex\u201d, Procedia Chemistry, Vol.19, pp:577-585. [7] Bensadoun. F, Depuydt. D, Baets. J, Verpoest. I, Van Vuure. A.W, 2017, \u201cLow velocity impact properties of flax composites\u201d, Composite Structures, Vol.176, pp:933-944. [8] Changlei Xia, Jason Yu, Sheldon Q. Shi, Ying Qiu, Liping Cai, Felix Wu. H, Han Ren, Xu Nie and Hualiang Zhang, 2017, \u201cNatural fibre and aluminum sheet hybrid composites for high electromagnetic interference shielding performance\u201d, Composites Part B, Vol.114, pp:121- 127. [9] Ming Liu, Andreas Baum, Jurgen Odermatt, Jens Berger, Liyun Yu, Birgitte Zeuner, Anders Thygesen, Jesper Holck and Anne S. Meyer, 2017, \u201cOxidation of lignin in hemp fibres by laccase: Effects on mechanical properties of hemp fibres and unidirectional fibre/epoxy composites\u201d, Composites: Part A, Vol.95, pp:377-387. [10] Bambach. M.R, 2018, \u201cGeometric optimisation and compression design of natural fibre composite structural channel sections\u201d, Composite Structures, Vol.185, pp:549-560. [11] Ravi Prakash Babu K., Rao P.K.V., Ali M.A., Raghu Kumar B., 2019, \u2018Crack simulations in shaft-hub-blade system using modal analysis\u2019, International Journal of Recent Technology and Engineering, 8(1), PP.2646-2650. [12] Bala Subramanyam P.N.V., Rao B.N., Prakash R.L., Rajavardhan D.S.K., Sankarudu B.S.K., Ganesh P.,2019, \u2018Modal analysis on go-kart chassis\u2019, International Journal of Engineering and Advanced Technology, 8(4), PP.1701-1705. [13] Syed K., Mannam H., Koya M.K., Meduri S.K., Laxmi V.S., Kasturi S.P., 2019, \u2018Dynamic simulation of gas turbine blade using finite element analysis\u2019, International Journal of Mechanical and Production Engineering Research and Development, 9(), PP.745-751. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 [14] Dama K.K., Surya S., Babu C., Pavan, Kiran R., 2018, \u2018Modal analysis of an automotive subsystem using cae techniques\u2019, International Journal of Mechanical Engineering and Technology, 9(3), PP. 1157-1162. [15] Deepak Kumar G., Kommuri N., Kota H., Prasanth V., Hameed S, 2018, \u2018Vibrational analysis of a human body on an automobile\u2019, Journal of Advanced Research in Dynamical and Control Systems ,10(6), PP. 39-45. [16] Devi S.K., 2018, \u2018The optimization of layer densities for multilayered insulation systems by using ansys\u2019, International Journal of Mechanical and Production Engineering Research and Development ,8(4), PP. 1087-1098. [17] Balasubramanyam P., Nageswara Rao B., Banerjee M., Lakshmana Swamy B., 2019, \u2018Impact analysis on Go-Kart chassis with variable speeds using ansys 19.0\u2019, International Journal of Engineering and Advanced Technology, 8(6), PP.2614-2620. [18] Kancheti N., Reddy Vemula A., Reddy Gudibandla G., Krishna H., Bala Subramanyam P.N.V. 2019, \u2018Modeling and analysis of wheel rim using ansys\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(8), PP.415-418. [19] Bharath E.S., NagaSree Harsha O., Bala Subramanyam P.N.V., Reddy G.S.S., 2019, \u2018Impact analysis on honey comb structured go-kart bumper using ansys R19.1\u2019, International Journal of Innovative Technology and Exploring Engineering, 8(6), PP.396-400. [20] Ravi Prakash Babu K., Gupta S.K., Phaneendra Sai Sri D., Gautam K., Kumar C.M. 2017,\u2019Significance of modal analysis to detect and diagnose misalignment fault in turbine rotor\u2019,Journal of Advanced Research in Dynamical and Control Systems,9(Special issue 14),PP.2637-2648. [21] Babu K.R.P., Prasanth K.F.V., Sai V.V., Kumar M.V.S., Kumar B.N.S., Chaitanya B.S. 2017,\u2019Effect of pre-twist on free vibration characteristics of a cantilever beam with experimental validation\u2019,International Journal of Mechanical Engineering and Technology,8(5),PP:389-399. [22] T. Srinivasan, CH.V.L.N. Ram Gopal, P. Sasidhar, P. Vineeth Kumar and Shaik Sohail, 2020, Effect of change of cross section and Frequency Analysis of Anti-Roll Bar Using FEA, IOP Conf. Series: Materials Science and Engineering 954 (2020) 012027 IOP Publishing, doi:10.1088/1757-899X/954/1/012027"
        ]
    },
    {
        "image_filename": "designv11_63_0003215_j.automatica.2021.109870-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003215_j.automatica.2021.109870-Figure2-1.png",
        "caption": "Fig. 2. Leader\u2013follower configuration.",
        "texts": [
            " The head position \u03befh(t) is expressed as a point that lies the distance \u03c1 along the vertical bisection of the wheel axis in front of the vehicle, i.e., \u03befh(t) = [xfh(t), yfh(t), \u03b8fh(t)]T . The head position \u03befh(t) is given by \u03befh(t) = \u03bef (t)+\u03c1[cos \u03b8f (t), sin \u03b8f (t), 0]T . Differentiating \u03befh(t) with respect to time gives a nominal system of the head position \u03be\u0307fh(t) = f (\u03befh(t), uf (t)) = [cos \u03b8f (t),\u2212\u03c1 sin \u03b8f (t); sin \u03b8f (t), \u03c1 cos \u03b8f (t); 0, 1] uf (t), where uf (t) = [\u03c5f (t), \u03c9f (t)]T . A leader\u2013follower tracking control problem is considered in a global frame as shown in Fig. 2. In Fig. 2, \u03bef (t) = [pTf (t), \u03b8f (t)] T is the state of the follower vehicle with pf (t) = [xf (t), yf (t)]T , \u03ber (t) = [pTr (t), \u03b8r (t)] T is the state of the leader vehicle is with pr (t) = [xr (t), yr (t)]T , the control signal ur (t) = [\u03c5r (t), \u03c9r (t)]T is given to the leader vehicle in advance such that a reference trajectory is obtained for the follower vehicle. Therefore, a nominal system on the leader vehicle is given as follows \u03be\u0307r (t) = f (\u03ber (t), ur (t)), ur (t) \u2208 U. There exists sideslip phenomenon for wheels of the vehicles during tracking and formation",
            " A perturbed head position kinematics of the follower vehicle is given as \u03be\u0307fh(t) = f (\u03befh(t), uf (t)) + n(t), t \u2265 t0, (2) where n(t) is the external disturbance with n(t) = [nx(t), ny(t), 0]T , t0 \u2208 R is an initial time, \u03befh(t) \u2208 Rn and uf (t) \u2208 Rm are the state and the input of the vehicle in the perturbed nonlinear nonholonomic system (2), respectively. Note that n(t) \u2208 Rn is an external disturbance which is bounded with \u2225n(t)\u2225 \u2264 \u03b7. Robustness of the perturbed nonlinear nonholonomic system (2) is guaranteed according to stability analysis for MPC. Constraints for the control input uf (t) and the state \u03befh(t) are shown as follows uf (t) \u2208 U \u2286 Rm, \u03befh(t) \u2208 X \u2286 Rn, \u2200t \u2265 t0, where U \u2286 Rm and X \u2286 Rn are two compact convex sets. In Fig. 2, the Frenet\u2013Serret frames rO and fO are constructed for the leader vehicles and the follower vehicles, respectively. That is, the Frenet\u2013Serret frames rO and fO are fixed on the vehicles. The tracking position for the head position of the follower vehicle is a virtual structure point dp = [dx, dy]T in the Frenet\u2013 Serret frame rO fixed on the leader vehicle. The tracking error is pef (t) = [xef (t), y e f (t)] T in the Frenet\u2013Serret frame fO fixed on the follower vehicle, i.e., pef (t) = R(\u2212\u03b8f (t))[xr (t) \u2212 xfh(t); yr (t) \u2212 yfh(t)] + R(\u03b8rf (t))[dx; dy] with \u03b8rf (t) = \u03b8r (t) \u2212 \u03b8f (t) and R(\u03b8 (t)) = [cos \u03b8 (t),\u2212 sin \u03b8 (t); sin \u03b8 (t), cos \u03b8 (t)], where R(\u00b7) is the rotation matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000456_0142331219897430-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000456_0142331219897430-Figure1-1.png",
        "caption": "Figure 1. Two-dimensional engagement geometry.",
        "texts": [
            " The paper is prepared as follows. First, the guidance strategy and missile dynamics are formulated. The design method for IGC law in presence of fin failure is discussed, as well. Next, a brief introduction to spherical unscented Kalman filter (SUKF) is provided. Then, the performance of the new IGC approach is evaluated through simulations. Finally, a summary and conclusion are presented. The model of IGC and missile dynamics are presented below. A two-dimensional geometry of planar interception is shown in Figure 1. In the polar coordinate system, the relative position of pursuer and target is given by R and l as follows (Hwang and Tahk, 2006) _R=VT cos (l gT ) VM cos (l gM ), _l= VT R sin (l gT )+ VM R sin (l gM ), \u00f01\u00de where l 2 R is the line of sight (LOS) angle, R 2 R. 0 is the range along the LOS, VM and VT 2 R are the speeds of missile and target, respectively. gM and gT 2 R are missile and target\u2019s flight path angles, respectively. R denotes the real number and R\u00f8 0 denotes the non-negative numbers. Suppose that the missile has constant speed within the endgame (Shima et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001535_j.jterra.2020.08.003-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001535_j.jterra.2020.08.003-Figure4-1.png",
        "caption": "Fig. 4. Real double wishbone suspension (a); Vi",
        "texts": [
            " Suspension Connection point X [m] Y [m] Z [m] Front suspension Lower arm \u2013 front connection 2.1755 0.223 0.103 Lower arm \u2013 rear connection 1.8335 0.223 0.103 Upper arm \u2013 front connection 2.1755 0.209 0.076 Upper arm \u2013 rear connection 1.8335 0.209 0.076 Front spring and shock absorber connection 2.0320 0.211 0.240 Rear suspension Trailing arm connection 0.3910 0.547 0.071 Trailing arm connection 0.3910 0.547 0.071 Rear spring and shock absorber connection 0.3640 0.545 0.500 The front suspension is a double wishbone (Fig. 4). The joints between the suspension arms and the chassis have bushings. The suspension arms have low mass and inertia relative to the sprung mass. Therefore, to reduce the number of equations, virtual stiff springs and virtual dampers are used, instead of bodies, to model the suspension arms. The virtual model stiff springs and virtual dampers are shown in black in Fig. 4b. The real suspension spring and shock absorber in the vehicle model are shown on Fig. 4b in gray color. The magnitude of the two systems are different. The vir- tual stiff spring stiffness is 5 105N/m and the virtual damping constant is 1000 Ns/m (it is assumed that this is the bushing stiffness and damping). The real suspension spring and shock absorber are defined as linear (k = 122625 [N/m], c = 1285 [Ns/m]). The force from the suspension spring and shock absorber are distributed among the lower arm connector points according to the distance between the connection points. The steering constraint forces on the steering wheel are neglected at this stage, because we are not interested in the torque on the steering wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002059_012021-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002059_012021-Figure1-1.png",
        "caption": "Figure 1. Structural (a) and hydraulic (b) diagrams of a forest fire ground-throwing machine with an energy-saving hydraulic drive. 1 - frame; 2 - linkage mechanism; rotor-thrower - 3; blades - 4, casingripper - 6; ploughshare - 7; windows - 8; safety knives - 9; support ski - 10; hydraulic motor - 11; distributor - 12; pressure hydraulic line - 13; hydroaccumulator - 14; check valve - 15; pump - 16; safety valves - 17 and 18.",
        "texts": [
            " Therefore, it is necessary to carry out additional theoretical and experimental studies of new designs of forest fire ground-throwing machines with an energy-saving hydraulic drive of active working bodies, which make it possible to improve the technology of laying mineralized strips and extinguishing forest ground fires. The aim of the study is to increase the efficiency of the working process of the soil-throwing machine by substantiating the parameters of the energy-saving hydraulic drive of the rotor-thrower driving the rotor. We have carried out theoretical studies aimed at developing a mathematical model for the joint work of passive and active working bodies with an energy-saving hydraulic drive of a forest fire groundthrowing machine, whose design is protected by a patent for a utility model (Figure 1) [8]. Mathematical model of the working process of the energy-saving hydraulic drive of the rotor- thrower, flowing in the hydraulic and mechanical subsystems:     2 0 2 1 1 EG EPN N \u041c \u041cR R \u041c R P i Gi i Pi \u0421\u0422 \u0412\u0422 R i i Q t P Pd f df J J J r F r F \u041c k dfdt dt dt                    (1) where JM, JR, JP \u2013 the moments of inertia of the hydraulic motor, rotor, transmission, reduced to the axis of rotation of the rotor; QM \u2013 hydraulic motor working fluid flow; P\u041c \u0438 P0 \u2013 pressure in the pressure head and drain hydraulic lines of the hydraulic motor; NEG \u0438 NEP \u2013 number of soil elements; ri \u2013 distance from the rotor axis to the i-th soil element or obstacle; F\u03c4Gi \u0438 F\u03c4Pi \u2013 tangential components of the force of action of the i-th soil element and obstacles on the rotor; \u041c\u0421\u0422 \u2013 the moment of the dry friction force during the rotation of the rotor; k\u0412\u0422 \u2013 coefficient of viscous friction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002260_cac51589.2020.9327221-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002260_cac51589.2020.9327221-Figure3-1.png",
        "caption": "Fig. 3. Dynamics Model",
        "texts": [
            " 1963 TABLE I MEANING OF SYMBOLS Symbol Parameter name a Length of tractor\u2019s centroid to tractor front axle b Length of tractor\u2019s centroid to tractor rear axle c Length of tractor\u2019s centroid to fifth wheel d Length from semi-trailer centroid to semi-trailer axle e Length of semi-trailer centroid to fifth wheel \u03b8 Articulation angle between tractor and semi-trailer \u03b4f Corner of front wheel x2 The vertical position of the semi-trailer\u2019s centroid y2 The horizontal position of the semi-trailer\u2019s centroid \u03c81, \u03c82 Yaw angle of tractor and semi-trailer \u03b21, \u03b22 Side slip angle of tractor and semi-trailer r1, r2 Yaw rate of tractor and semi-trailer u1, u2 Speed of tractor and semi-trailer m1,m2 Mass of tractor and semi-trailer I1zz , I2zz Moment of inertia of tractor and semi-trailer around stern axis F1, F2 Lateral forces of the front and rear axles of the tractor F3 Lateral forces of the axle of the semi-trailer F4 Simplified force between tractor and semi-trailer k1, k2, k3 Tractor front, rear axle and semi-trailer axle lateral stiffness Based on the principle of vehicle dynamics, a three-degreeof-freedom dynamic model is established shown in Fig.3. The meaning of each variable in the picture is also shown in Table I. In Fig.3, the tractor and semi-trailer are simplified as rods with mass of m1 and m2 respectively. When the road curvature is kept in a small range, the articulation angle \u03b8 between the tractor and the semi-trailer will also be kept in a small range. Ignoring the influence of aerodynamics and road slope angle, the dynamic equation of tractor can be written as follows according to Newton\u2019s law when the longitudinal speed is constant: { m1u1(\u03b2\u03071 + \u03c8\u03071) = F1 + F2 + F4 I1zz r\u03071 = F1a\u2212 F2b\u2212 F4c (6) According to Newton\u2019s law, the dynamic equation of the trailer can be obtained as follows:{ m2u2(\u03b2\u03072 + \u03c8\u03072) = F3 \u2212 F4 I2zz r\u03072 = \u2212F3d\u2212 F4e (7) Assuming the tire always working within linear region, the tractor front axle tire force F1, the tractor rear axle tire force F2 and the trailer rear axle tire force F3 can be calculated by the following formula respectively:\u23a7\u23a8\u23a9 F1 = k1(\u03b21 + a r1 u1 \u2212 \u03b4f ) F2 = k2(\u03b21 \u2212 b r1 u1 ) F3 = k3(\u03b22 \u2212 d r2 u2 ) (8) Since the tractor and semi-trailer are connected by hinges, there exists the following kinematic constraints between the tractor and semi-trailer: \u03b2\u03071 \u2212 \u03b2\u03072 \u2212 c u1 r\u03071 \u2212 e u2 r\u03072 + r1 \u2212 r2 = 0 (9) According the Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002723_icitamee50454.2020.9398415-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002723_icitamee50454.2020.9398415-Figure5-1.png",
        "caption": "Fig. 5. Wheelchair's design",
        "texts": [
            " The ADXL335 Accelerometer circuit consists of some 5-pin switches; the first pin is connected to 5v; the second pin is connected to the A2 Arduino port; the third and fourth pins are connected to the Arduino A3 port, and the fifth pin is connected to the Ground [23]. D. Tools Designing In designing tools on electric wheelchairs, the authors use the following tools and materials: ADXL335 accelerometer, wheelchair, battery, Arduino minimum system, motor, and DC motor. The Illustration of tool design can be seen in Fig. 5. III. RESULT AND DISCUSSION Testing and measurement of the electric wheelchair with Accelometer was done through several steps. Table I is the result of an accelerometer command testing on an electric wheelchair with orders; forward, backward, turn left, and turn right. Testing is done 25 times. Then, the results of the experiments carried out are calculated using the accuracy formula. The following results were obtained: 88% forward, 84% turn right, 88% turn left, 92% reverse done accurately. The error analysis regarding this test is; the user is not accustomed to using the head as a controller so that the movement is not constant and the range of head movement reading programs between one command and another command is too close"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000278_wisee.2019.8920360-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000278_wisee.2019.8920360-Figure1-1.png",
        "caption": "Fig. 1. A two-wheeled mobile robot",
        "texts": [
            " The neural network will be trained online and once convergence is achieved with a desired precision, it can be used for the prediction of actions according to the current state of the UAV. Algorithm 2 presents Q-learning procedure for a UAV based on the method in [13]. Consider now three mobile robots modeled as two-wheeled vehicles. In this section, the dynamics of these unmanned ground vehicles (UGV) and the way they track a trajectory specified by a UAV is illustrated. The kinematic model variables of a two-wheeled mobile robot are shown in Fig 1. The trajectory of the robot is controlled by properly adjusting the angular velocities of the right and left wheels. Here, (x, y) represents the generalized coordinates of the center of mass of the robot and \u03b8 denotes its heading w.r.t. the x axis. The linear and angular velocities of the robot are expressed by v and \u03c9, respectively. Assuming there is no slippage on the surface, the kinematic model is described by [14] :x\u0307y\u0307 \u03b8\u0307  = cos \u03b8 0 sin \u03b8 1 0 1 (v \u03c9 ) (7) Let the reference trajectory for tracking be qr = [xr, yr, \u03b8r], and the current state of the robot be qc = [xc, yc, \u03b8c]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002325_cac51589.2020.9327853-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002325_cac51589.2020.9327853-Figure1-1.png",
        "caption": "Fig. 1. Diagram of a two-link robotic manipulator.",
        "texts": [
            " Simulation results shown in Section IV demonstrate the feasibility and efficiency of the proposed strategy. Lastly, section V gives the conclusions of this paper. 20 20 C hi ne se A ut om at io n C on gr es s ( C A C ) | 9 78 -1 -7 28 1- 76 87 -1 /2 0/ $3 1. 00 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ C A C 51 58 9. 20 20 Authorized licensed use limited to: City&#44; University of London. Downloaded on May 17,2021 at 04:05:06 UTC from IEEE Xplore. Restrictions apply. 2688 II. PROBLEM FORMULATION The model of a robotic manipulator system composed of two rigid links (see Fig. 1) considered in this paper is given as follows M ( ) \u0308+ C ( , \u0307 ) \u0307+G ( ) = \u03c4 (1) where the matrix M ( ) \u2208 R2\u00d72 is symmetric and positive, C ( , \u0307 ) \u2208 R2\u00d72 is a matrix represents the CoriolisCentrifugal forces, while G ( ) \u2208 R2 represents the gravity vector. \u2208 R2 represents the angular position, while its first and second order time derivative \u0307, \u0308 are the velocity and acceleration of angular, respectively. In addition, the input torque vector is denoted by \u03c4 \u2208 R2. It should be stressed that the studied system has the following properties: Property 1: The inertial matrix is denoted by M ( ), which is bounded"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure51.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure51.3-1.png",
        "caption": "Fig. 51.3 Kinematic model of multi-DoF instrument with the novel end-effector",
        "texts": [
            " \u2022 Link length (ai): The length of the common normal, which is the distance between the previous z-axis (zi\u22121-axis) and the current z-axis (zi -axis). \u2022 Link twist angle (i: The angle around the common normal between the previous z-axis (zi\u22121-axis) and current z-axis (zi -axis). \u2022 Link offset (di): The distance between the previous x-axis (xi\u22121-axis) and the current x-axis (xi -axis), along zi\u22121-axis. \u2022 Joint angle ( i): The angle around the z-axis between the previous x-axis (xi\u22121axis) and the current x-axis (xi -axis). The kinematic model of the multi-degree of freedom instrument with a novel end-effector mechanism is proposed in Fig. 51.3. Based on the kinematic model, the homogeneous transformationmatrix for each joint is calculated. TheD-H parameters for each joint in Fig. 51.3 are summarized in Table 51.2. Table 51.1 summarizes the range of the variable parameters of each joint for this novel end-effector mechanism. The transformationmatrix between every two successive frames canbe formulated using the homogeneous transformation matrix equation, and the D-H parameters summarized in Table 51.2 for each joint. The transformation matrices for joint 1 to joint 7 are calculated below based on the values of different joint angles and link lengths (Table 51.2). 0T1= \u23a1 \u23a2\u23a2\u23a3 c1 \u2212s1 0 0 s1 c1 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6; 1T2= \u23a1 \u23a2\u23a2\u23a3 c2 0 s2 0 s2 0 \u2212c2 0 0 1 0 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6; where cn : cos(\u03b8n), sn : sin(\u03b8n), sn\u22121n : sin(\u03b8n\u22121 + \u03b8n), sn n\u22121 : sin(\u03b8n\u22121\u2212\u03b8n), cn\u22121n : cos(\u03b8n\u22121 + \u03b8n), cn n\u22121 : cos(\u03b8n\u22121 \u2212 \u03b8n) Now, the position of the end-effector tip with respect to the fixed reference frame is given by: BTW =0 T7 =0 T1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000295_icems.2019.8921961-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000295_icems.2019.8921961-Figure4-1.png",
        "caption": "Fig. 4. Temperature Cloud of Hub Motor and Cooling Water at Water Velocity of 0.2m/s",
        "texts": [
            " In order to determine the optimal flow velocity of watercooled motor, it is necessary to compare the maximum temperature of each part of the motor under different flow velocities. The inlet boundary condition of cooling water is velocity-inlet, the initial temperature of cooling water is 55\u00b0C, and the outlet boundary condition of cooling water is pressure-outlet. The flow velocity is set to 0.1m/s, 0.2m/s and... 3m/s, respectively. When the water velocity is 0.2m/s, the steady state cloud and cooling water temperature of the whole temperature field of the motor are shown in Fig. 4. According to Fig. 4, it is observed that the inlet temperature is 54.85\u00b0C, the outlet temperature is 95.78\u00b0C, the temperature difference between the inlet and outlet reaches 40.93\u00b0C, and the temperature rise of cooling water is obvious, which verifies the effectiveness of cooling water channel. The maximum temperature of each part of the motor under the condition of cooling water speed of 0.2m/s and natural cooling is compared as shown in Table 7. In Figure 4, it is observed that the overall cooling effect of the motor is very obvious. The maximum temperature of the winding is 140.21\u00b0C, which is 43.98\u00b0C different from the maximum value of the natural cooling winding, and the maximum insulation temperature is 140.1\u00b0C, which is 43.1\u00b0C different from the maximum value of the natural cooling insulation. Because the cooling water in the stator yoke and waterway can exchange heat well, the temperature of the yoke is lower than that of the teeth, so the temperature of the winding and insulation close to the stator yoke is lower than that close to the teeth of the stator",
            " This is because the temperature of the cooling water at the inlet of the channel is lower than that at the outlet, and the temperature of the stator decreases unevenly. After heat exchange, the temperature of the winding must also be uneven. At the same time, the highest temperature of winding is also lower than the performance reference temperature of class H insulation, and the cooling effect is very obvious. Because the cooling water channel is close to the stator core, the cooling effect of the stator is the most remarkable. In Figure 4, the maximum stator temperature is 124.55\u00b0C, which is 45.87\u00b0C lower than that of natural cooling. Because the cooling water temperature at the inlet of the water channel is the lowest, the temperature of the cooling water increases gradually after heat exchange with the stator, so the temperature distribution on the inner surface of the stator yoke is uneven, and the cooling effect of the stator at the inlet of the cooling water is the best. Similar to natural cooling, the temperature of stator tip is lower, because the stator tip is located in the part with faster air flow rate and can exchange heat with air, so the temperature of this part is relatively low. The central part of the winding in the stator slot heats most seriously, which results in higher temperature in the central part of the core near the winding slot. Compared with stator and winding, the cooling effect of permanent magnet and rotor is not obvious, but the temperature also decreases. The maximum temperature of permanent magnet decreases by 5.61\u00b0Cand that of rotor by 5.43\u00b0C. The distribution of temperature nephogram is basically the same as that of natural cooling. From Figure 4, it can be seen that the cooling effect of the cooling water channel on the stator and winding parts of the motor is relatively obvious. Combining with other water speed simulation results, the maximum temperature and cooling water speed curves of each part of the motor are obtained, as shown in Figure 5. From the maximum temperature curve of each part of hub motor at different water speeds in Fig. 5, it is observed that the temperature of stator, winding and insulation decreases most significantly in the range of 0-0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000608_s00170-020-05023-4-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000608_s00170-020-05023-4-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the S-VTS model",
        "texts": [
            " The fractal pattern always causes interior defects at overlapping area and increases the planning workload. Therefore, we choose the raster pattern to control the scanning path and further study the deposition strategies for height difference. Based on the geometric characteristics of the curved surface structure, freeform surface is adopted to slice the 3D entity by S-VTSmodel. According to the CADmodel of the curved surface structure, a NURBS surface model Mtop is extracted to obtain the two-dimensional height differences\u0394HX and \u0394HY, as shown in Fig. 2. Here the scanning direction is defined asYaxis (if\u0394HX \u2264\u0394HY), and the overlapping direction is X axis because more slant surface can be achieved by changing the scanning speed. Variable thickness layers L1~LN will be sliced by using multi-surface (blue surfaces in Fig. 2). According to the stagnation points on the intersection curve, each layer is divided into M portions and each portion i(1, 2,\u22ef,M) is sliced individually. The height difference of every portion in both directions should be within the optimal height range. To ensure effective deposition of each layer, the number of slicing surface Ni in each portion should satisfy the following constraints: HSTi=Ni\u2264hmax HSBi=Ni\u2265hmin \u00f01\u00de where HSTi and HSBi are the lowest and highest height of ith portion; hmin and hmax are the optimal height range"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001154_rpj-09-2018-0243-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001154_rpj-09-2018-0243-Figure2-1.png",
        "caption": "Figure 2 Design of specimen for torsion testing of IN718: a) geometric dimensions of specimen in inches, b) experimental setup for torsional fatigue testing (note: specimen was not subjected to axial loading)",
        "texts": [
            "58546 Lambert (2016) SLM system/vertical Surface roughness effect on HCF stress-life/R = 0.1/ HIP and as-built/1200\u00b0F and 800\u00b0F Konecna et al. (2016) Renishaw SLM 250/horizontal and vertical Fatigue crack behavior/HCF through plane bending tests at R = 0/as-built surface/room temperature Yadollahi et al. (2018) Concept Laser M1/vertical and horizontal Stress intensity factor range against crack growth rate relation/HCF force-control, R = 0.1 and R = 1/as-built and machined surface conditions/room temperature Rapid Prototyping Journal developed is as presented in Figure 2a. To serve as a comparison, a \u00bd in diameter wrought annealed IN718 rod was machined based upon the geometric design. The wrought annealed specimen was heat-treated based upon the ASTM Standard Designation B637-18 and followed the specific heat-treatment type: \u201csolution annealing (924 to 1,010\u00b0C), hold \u00bd h (commensurate with cross-sectional thickness), cool at rate equivalent to air cool or faster (ASTM, 2018).\u201d Experimental testing on both additively manufactured and wrought IN718 specimens were then performed",
            " Gordon Rapid Prototyping Journal DMLS IN718 specimens manufactured along each build orientation were subject to room temperature torsional fatigue tests, under the same experimental conditions as its conventionally manufactured counterpart. Completely reversed (Rf = 1) torsional fatigue tests were performed at an angle of twist cycling range of Df = 615\u00b0 and twisting rate of 1.654deg/sec, using the MTS EM Bionix torsion test system, with a maximum torque of 50N-m. The experimental setup with the test device is as shown in Figure 2b. Data acquisition rate was set at 10Hz and the axial load was kept at 0N, for the purpose of this study. Data was outputted as torque versus angle of twist, which were converted to shear stress and shear strain using the following relations. t \u00bc Td 2J (1) The output torque,T, was used to determine the shear stress, t , from equation (1), using the inner gauge diameter, d, and polar moment of inertia, J, based upon the specimen geometry presented in Figure 2a. g \u00bc f d 2L (2) The shear strain, g , was determined from equation (2), using the inner gauge diameter, d, inner gauge length, L, and angle of twist, f . J \u00bc pd4 32 (3) The polar moment of inertia, J, was determined using equation (3) and is a function of the inner gauge diameter, d. 2.4Metallurgical analysis preparation To relate microstructural characteristics (i.e. grain size) to reported torsional fatigue response, IN718 specimens of X, Z and XZ45\u00b0 build orientation were sectioned using a slow-speed Isomet cutter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000498_ab6d4e-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000498_ab6d4e-Figure1-1.png",
        "caption": "Fig. 1: General illustration of laser-powder bed fusion process showing different regions and phenomena likely to be present in different working modes of LPBF process.",
        "texts": [
            "[13\u201322] However, experimental observations evidenced that in some particular ranges of power and scanning speed, during powder consolidation and melt pool formation, phenomena like splattering from the melt pool, ejection of powder particle, and extreme key hole formation can become significant.[9,10,23\u201330] Simulation models with improved fidelity would then include physics like wetting and consolidation of powder, handling of the movement of multiple phases, irradiance of hot surface, recoil pressure onto the melt pool surface and the interaction of laser beam and the melt pool.[8,16,31,32] Fig. 1 gives a general illustration of the physics encountered in different working modes of LPBF process. 2 A cc ep te d M a us cr ip t The starting point of simulation model for SLM process is laser energy absorption. Laser absorption in SLM is modeled in various ways. However, experimental observations of the time required for metal powder consolidation,[18,33] for relatively fast and sharp laser beam commonly used in SLM today, confirm that laser absorption is likely to happen in powder-bed before consolidation takes place"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000124_012164-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000124_012164-Figure3-1.png",
        "caption": "Figure 3. Frame modeling uses 2015 Autodesk inventor professional software.",
        "texts": [],
        "surrounding_texts": [
            "The research was carried out in the Computer Laboratory and Design of the Mechanical Engineering Department of Musamus University. The research method used was an experimental method (computer simulation). The strength of construction with varied loads using the 2015 Autodesk inventor professional software. ICROEST IOP Conf. Series: Earth and Environmental Science 343 (2019) 012164 IOP Publishing doi:10.1088/1755-1315/343/1/012164 Free variables: a load of raw materials (nuts) and machine components. Non-independent variables : main stress distribution, displacement, and safety factor. To determine the strength of frame construction loading is given on the Y-axis of 500 N, 600 N, 700 N, 800 N, 900 N, and 1000 N ICROEST IOP Conf. Series: Earth and Environmental Science 343 (2019) 012164 IOP Publishing doi:10.1088/1755-1315/343/1/012164"
        ]
    },
    {
        "image_filename": "designv11_63_0002755_tmag.2021.3074935-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002755_tmag.2021.3074935-Figure1-1.png",
        "caption": "Fig. 1. Four slot/pole combinations and their coil connections of DCB-VRMs. Nr /Pdc/Pa are: (a) Model 1: 8/6/2. (b) Model 2: 5/3/2. (c) Model 3: 7/3/2. (d) Model 4: 7/6/1.",
        "texts": [
            " Date of publication April 22, 2021; date of current version June 23, 2021. Corresponding author: S. Jia (e-mail: shaofengjia@ xjtu.edu.cn). Color versions of one or more figures in this article are available at https://doi.org/10.1109/TMAG.2021.3074935. Digital Object Identifier 10.1109/TMAG.2021.3074935 The design of topologies is still in accordance with the principle of magnetic field modulation [4]. Possible slot/pole combinations, coil connections, and current configuration of some 12-stator-slot machines having distributed windings are shown in Fig. 1, based on the following criteria for the choice of slot/pole combinations: Pa = |Pdc \u2212 Nr \u2213 k Ns| (1) where Pa and Pdc are the armature pole pair number and dc MMF pole pair number, respectively, and Ns and Nr are the stator and rotor slots, respectively. k is 0 or 1. When k is 1, it represents the influence of introduced stator slotting effect. 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index",
            " Winding factor is an important parameter to evaluate the utilization of the coil. It can be calculated that winding factor of Models 1, 2, and 3 is 1 and winding factor of Model 4 is 0.966. Some main parameters of the proposed dc-biased VRMs having distributed windings are shown in Table I, which are the same as the parameters in the previous article [7], in order to compare the torque density with different windings fairly. The direction of dc current injected into each winding has been shown in Fig. 1. It can be seen that Pdc is 3 if the direction of the injected dc currents changes alternately for every two stator slots while Pdc is 6 if the direction of the injected dc currents changes alternately for every one stator slot. For Model 1 in Fig. 1(a), the phase currents are iA = Idc + \u221a 2Iac sin(\u03c9et + \u03b1) iB = Idc + \u221a 2Iac sin(\u03c9et + \u03b1 \u2212 2\u03c0/3) iC = Idc + \u221a 2Iac sin(\u03c9et + \u03b1 + 2\u03c0/3). (2) While for Models 2\u20134 in Fig. 2(b)\u2013(d), the phase currents are i A+ = \u221a 2Iac sin(\u03c9et + \u03b1)+ Idc i A\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1)\u2212 Idc iB+ = \u221a 2Iac sin(\u03c9et + \u03b1 \u2212 2\u03c0/3)+ Idc iB\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1 \u2212 2\u03c0/3)\u2212 Idc iC+ = \u221a 2Iac sin(\u03c9et + \u03b1 + 2\u03c0/3)+ Idc iC\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1 + 2\u03c0/3)\u2212 Idc (3) where Idc is the dc component, Iac is the ac component, and \u03b1 is the phase A current angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002423_012128-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002423_012128-Figure5-1.png",
        "caption": "Figure 5. Boundary conditions for analysis.",
        "texts": [
            " Series: Materials Science and Engineering 1070 (2021) 012128 IOP Publishing doi:10.1088/1757-899X/1070/1/012128 These parameters are assigned 3 levels to determine the optimum combination (Table 1). The desired functionality of the unit cell is to provide the maximum negative Poisson\u2019s ratio with minimal induced stress, while subjected to an external force. L9 orthogonal array is used to analyse feasible test cases (Table 2). A compressive load of 25 N is applied in the axial direction of the cubic unit cell with its bottom link fixed (Figure 5). Tests are conducted with the combinations shown in the orthogonal array to identify the combination that provides the desired output. The optimal combination of the factors and levels were determined through analysis. Not all test cases could be analysed, as test cases 1,2,4 and 7 rendered an infeasible assembly for the combination of link length and link angle. A link length of 25mm and a link angle of 70\u00b0 provides the favorable output with a Poisson ratio of -0.576 and an induced stress of 228"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure12-1.png",
        "caption": "Figure 12. 3D model of the caisson.",
        "texts": [
            " It results that, at the same load, the deformation in the case of the composite door would be about eight times higher than in the case of the metal door. So, indeed, the composite door is significantly more elastic than the metal one. It was concluded that the surface of the door panel should be stiffened. The stiffening can be done by making a caisson, on the inside of the door panel, made of composite materials, as shown in Figure 11 (a). The caisson is made using a corrugated glass fiber marker, glued to the inside of the door face, over which another smooth layer of fiberglass is glued as seen in detail in Figure 11(b). Figure 12 shows the 3D model of the corrugated part and the inner panel of the caisson. Following the modal analysis made on the 3D model of the door with the stiffened surface, other values were obtained for the first eigenvalues of vibrations, previously determined. Figures 13 to 15 present comparatively the two variants of doors, i.e. the simple and the reinforced version, for the six determined eigenmodes. In the Appendix are presented the eigenmodes 7\u201310 (Figures 21 and 22). Table 1 presents, for comparison, the eigenvalue modes and the eigenmodes for the two versions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002448_physreve.103.033001-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002448_physreve.103.033001-Figure1-1.png",
        "caption": "FIG. 1. Schematic overview of the system. Thin sheet with total length L and bending modulus B is compressed symmetrically between the two sides of a rectangular closed chamber. The dimensions of the chamber are Lx \u00d7 Ly. A Cartesian coordinate system is located on the left edge of the sheet at a height Ly/2 above the bottom wall. In the initial setup the valve is closed and the two sides of the chamber are filled with an incompressible fluid with equal volumes, Vu = Vd. Under this setup the sheet accommodates an elastic shape (solid-gray line) that is higher in energy compared with the first mode of buckling (dashed-gray line). Then we open the valve and allow a controllable exchange of fluid between the two sides of the chamber. The sheet now spontaneously relaxes its bending energy until it reaches the lower energetic state.",
        "texts": [
            " These morphological transitions release elastic energy that, if properly managed, can be exploited for the purpose of energy harvesting; see recent reviews on the subject in Refs. [20\u201329] and references therein. Herein, we study a mechanical system that converts the elastic energy stored in a thin sheet, into a hydrodynamic pressure that potentially can enable a flow. Our system consists of an inextensible sheet that is laterally confined between the two sides of a rectangular closed chamber (Fig. 1). Had we kept the system in this configuration the sheet would buckle to accommodate the external compression, which is manifested by the first mode of buckling in the framework of Euler\u2019s elastica [30,31]. However, we further fill the chamber with an incompressible fluid that fixes the volumes above and below the sheet. This additional confinement requires the sheet to accommodate a configuration that is higher in energy compared with the trivial first mode of buckling. Then, we connect the two sides of the chamber and allow an exchange of fluid between them, i",
            " III we derive an approximated solution to the asymmetric and the symmetric branches in the limit of a small lateral confinement and analyze the transition between them. In Sec. IV we solve our model numerically beyond the small amplitude approximation and discuss the effect of the lateral compression on the maximum pressure drop in the chamber. In Sec. V we conclude and discuss possible extensions for future studies. II. FORMULATION OF THE PROBLEM An inextensible thin sheet with bending modulus B, thickness t , and total length L is confined between the two sides of a rectangular closed chamber, see Fig. 1. The horizontal, the vertical, and the width dimensions of the chamber are denoted by Lx (L Lx), Ly, and W , respectively. The two sides of the chamber, above and below the elastic sheet, are filled with an incompressible fluid with volumes Vu and Vd, respectively. Initially, we set Vu = Vd. Then we connect the two sides of the chamber and allow a controllable exchange of fluid between them, see Fig. 1. Since under this initial setup the elastic configuration does not accommodate its lowest energetic state, we would expect it to snap and enhance the fluid exchange. While this change in the sheet\u2019s configuration can result in a rapid and irreversible flow, in this work we neglect dynamic effects and focus on the slow, quasistatic, evolution of the system. The amount of fluid transferred between the two sides of the chamber is our control parameter. An alternative scenario in which the pressure difference in the chamber is the control parameter is discussed in Appendix A",
            " Hopefully, the answer to these fundamental questions will lay the groundwork for advanced technological applications that exploit the snap-through instability to the design of mechanical switches [19,60\u201362], or memcapacitors in electrical circuits [63,64]. ACKNOWLEDGMENTS We thank Benny Davidovitch, Lior Atia, and Yuri Feldman for stimulating discussions. This work was supported, in part, by the Pearlstone Center for Aeronautical Studies. APPENDIX A: EVOLUTION OF A PRESSURE-CONTROLLED SYSTEM In this Appendix we examine the evolution of the system when the pressure difference, instead of the volume difference, is the control parameter. To do that, we slightly modify the schematic overview of the system, Fig. 1, and consider a chamber that is endowed with two pistons at its upper and bottom walls. These pistons fix the upper and lower pressures, Pu and Pd, respectively, as seen in Fig. 9. Initially we set Pu = Pd (Pud = 0), and then we increase the upper pressure while keeping the lower pressure fixed, i.e., Pud 0. Given the properties of the sheet, L and B, the dimensions of the 033001-10 chamber Lx \u00d7 Ly, and the pressure difference Pud, we wish to find the spatial orientation of the sheet. Note that, different from the main text, in this Appendix the shape changes of the sheet are not a result of a spontaneous process but they are driven by the external pressures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure17-1.png",
        "caption": "Fig. 17. Dumpable structure",
        "texts": [],
        "surrounding_texts": [
            "The dumpable structure is to pour the bird thorn to one side after the anti-bird thorn is folded, which provides more work space for maintenance. The anti-bird spur and the base part are connected by screws. When the anti-bird device needs to be tipped over, the screw of the connection part can be loosened with a screw wrench to tilt the bird spur to one side to avoid affecting operation and maintenance. After the overhaul is over, the bird spines will be straightened and fixed, and then the bird spines can be opened."
        ]
    },
    {
        "image_filename": "designv11_63_0000133_s11071-019-05325-7-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000133_s11071-019-05325-7-Figure1-1.png",
        "caption": "Fig. 1 A Chaplygin sleigh with a free rotor",
        "texts": [
            " One of them is the carrying body\u2014a platform (Chaplygin sleigh) which slides on a horizontal plane. The point P fixed on the platform cannot slide in some direction n fixed relative to the platform: (vP , n) = 0, (1) where vP is the velocity of point P . The constraint (1) can be realized by means of the knife edge or the wheel pair, inwhich there is no slipping at the points of contact of the wheels with the plane [2]. The other body is an unbalanced rotor. It is fixed on the platform at some point R and rotates freely in the horizontal plane (see Fig. 1). We define two coordinate systems: \u2013 an inertial (fixed) coordinate system Oxy; \u2013 a noninertial coordinate system O1x1x2 attached to the platform, with origin O1 at the intersection of the perpendicular to the plane of the knife edge at the point of contact P with the straight line parallel to the plane of the knife edge and passing through the center of mass of the platform C . Let us introduce generalized coordinates. Let r = (x, y) be the radius vector of point O1 in the fixed coordinate system Oxy",
            " We represent the kinetic energy of the sleigh in the form Ts = 1 2 ms(x\u0307 2 + y\u03072) \u2212 msc(x\u0307 sin\u03c8 \u2212 y\u0307 cos\u03c8)\u03c8\u0307 +1 2 (Is + msc 2)\u03c8\u03072, where ms and Is are, respectively, the mass and the moment of inertia of the sleigh and c is the distance from the center of mass of the sleigh to origin O1. The kinetic energy of the unbalanced rotor has the form Tr = 1 2 mrv 2 G + 1 2 Ir(\u03c8\u0307 + \u03d5\u0307)2, vG = r\u0307 + (\u03d5\u0307 + \u03c8\u0307)Jz s + \u03c8\u0307Jz\u03c1, s = (s cos(\u03d5 + \u03c8), s sin(\u03d5 + \u03c8)), \u03c1 = a\u03c4 + bn, Jz = ( 0 \u22121 1 0 ) , wheremr and Ir are the mass and the moment of inertia of the rotor, respectively, and distances s, a and b characterize the position of the center of mass of the rotor on the platform (see Fig. 1). Substituting (2), we obtain the kinetic energy in the form T\u0303 = mr + ms 2 (v21 + v22) + 1 2 Js\u03c9 2 s + 1 2 Jr\u03c9 2 r +mrs(a cos\u03d5 + b sin \u03d5)\u03c9s\u03c9r +(mra + msc)v2\u03c9s \u2212mrs(v1 sin \u03d5 \u2212 v2 cos\u03d5)\u03c9r \u2212 mrbv1\u03c9s, Js = Is + msc 2 + mr(a 2 + b2), Jr = Ir + mrs 2. (4) Wenote that, if the quasi-velocities are chosen in this way, the kinetic energy of the system does not depend on \u03c8 . Since the kinetic energy does not depend on the configuration variables\u03c8 , x and y, the equations ofmotion of a nonholonomic system can be represented in the following general form (for details, see [11]): d dt ( \u2202 T\u0303 \u2202v1 ) \u2212 \u03c9s \u2202 T\u0303 \u2202v2 = \u03bb \u2202Q \u2202v1 , d dt ( \u2202 T\u0303 \u2202v2 ) + \u03c9s \u2202 T\u0303 \u2202v1 = \u03bb \u2202Q \u2202v2 , d dt ( \u2202 T\u0303 \u2202\u03c9s ) + \u2202 T\u0303 \u2202\u03d5 \u2212 v2 \u2202 T\u0303 \u2202v1 + v1 \u2202 T\u0303 \u2202v2 = \u03bb \u2202Q \u2202\u03c9s , d dt ( \u2202 T\u0303 \u2202\u03c9r ) \u2212 \u2202 T\u0303 \u2202\u03d5 = \u03bb \u2202Q \u2202\u03c9r , (5) where \u03bb is the undetermined multiplier",
            " In addition, if the parameters of the sleigh and the rotor satisfy the condition |\u03b4| |mss|, \u03b4 = mra + msc, then the reduced system possesses two one-parameter equilibriumpoints atwhich the sleighmoves in a circle: \u03a3 \u2032 2 = { v1 = \u03a9 \u03b4 ( b\u03b4 + a \u221a m2 r s 2 \u2212 \u03b42 ) , \u03c9s = \u03a9, \u03c9r = \u03a9,\u03d5 = \u2212 arctan \u221a m2 r s 2 \u2212 \u03b42 \u03b4 } , \u03a3 \u2032 3 = { v1 = \u03a9 \u03b4 ( b\u03b4 \u2212 a \u221a m2 r s 2 \u2212 \u03b42 ) , \u03c9s = \u03a9, \u03c9r = \u03a9,\u03d5 = arctan \u221a m2 r s 2 \u2212 \u03b42 \u03b4 } , \u03a9 = const. The equilibriumpoints\u03a3 \u2032 2 and\u03a3 \u2032 3 correspond to a position of the rotor inwhich the center ofmass of the entire system (sleigh+rotor) is balanced relative to the knife edge (i.e., it lies on the axis O1x1). Indeed, one can show that for these equilibrium points the following relation holds: (msrC + mrrG , n) = 0, where rC and rG are the radius vectors of O1C and O1G, respectively (see Fig. 1). On the fixed level set of the energy integral M3 h , the equilibrium points \u03a3 \u2032 2 and \u03a3 \u2032 3 define isolated fixed points. For these equilibrium points, the characteristic polynomial has the form P(\u03bb) = p3\u03bb 3 + p2\u03bb 2 + p1\u03bb + p0. Explicit expressions for the coefficients pi , i = 0, . . . , 3 are rather cumbersome, but they can be obtained using any system of analytical calculations, for example, Maple or Mathematica. As we see, the characteristic polynomial is not symmetric. Consequently, the equilibrium points \u03a3 \u2032 2 and \u03a3 \u2032 3 can in the general case be asymptotically stable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000218_isncc.2019.8909197-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000218_isncc.2019.8909197-Figure3-1.png",
        "caption": "Fig. 3. Sketch structure and fabrication images of controllable polarity doublegate SiNWFETs from [12].",
        "texts": [
            " Following to its geometry and physics, the final electric way connection in the nanorelay is biconditional on the terminal values[15]. Even though they are based on different technologies, all the devices in [12]\u2013[15] have the same common logic abstraction, depicted by Fig. 2. In this work, we focus on double-gate controllable polarity SiNWFETs [12] to showcase the impact of novel logic representation forms in emerging technology synthesis. A device sketch and fabrication views from [12] are reported in Fig. 3 for the sake of clarity. While having an enhanced functionality, this device also presents an implementation overhead deriving from the second gate. Without a dedicated logic abstraction and synthesis methodology, the full potential of this technology may remain obscured by the extra fabrication cost. We propose in this paper a novel logic representation form, based on the biconditional connective, that naturally harnesses the operation of controllable polarity switches. By using this logic abstraction in synthesis, the full logic expressive power of double-gate controllable polarity SiNWFETs, but also of devices in [13]\u2013[15], is unlocked"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001028_physrevfluids.5.054003-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001028_physrevfluids.5.054003-Figure1-1.png",
        "caption": "FIG. 1. Schematic of a chambered doctoring system in a gravure process: liquid is filled into cavities on an engraved roller and the excess liquid is wiped off by a doctor blade [5].",
        "texts": [
            " For both model problems, it is found that increasing cavity width and wettability, decreasing wall steepness, or lowering the capillary number generally improves filling by allowing the contact line on the cavity to slip more. Shear thinning enhances contact-line motion via reduced viscosities near the dynamic contact line and, as a consequence, leads to improved cavity filling. DOI: 10.1103/PhysRevFluids.5.054003 I. INTRODUCTION Liquid filling of recessed features having micro- to nanometer scales plays a key role in various settings such as imprint lithography [1\u20134] and gravure printing (Fig. 1) [5]. Incomplete filling could lead to entrapment of air, which may cause bubble-type defects in printed/imprinted structures, compromising device functionality and reproducibility [6\u201311]. Feature filling is also highly relevant to other applications including microfluidics [12\u201314], lubricant-impregnated surfaces [15\u201317], and porous-media flow [18\u201321]. Whether recessed features are completely filled depends on the movement of three-phase contact lines and the deformation of the liquid-air interface [22,23]",
            " This configuration serves as a simple model of the filling process in imprint lithography [31] and is relevant to microfluidic [13,14] and porous-media flows [18\u201320]. In the second configuration [Fig. 2(b)], the flat plate moves horizontally above a stationary cavity and the flow is driven by the plate motion and a pressure gradient. This configuration serves as a 2469-990X/2020/5(5)/054003(19) 054003-1 \u00a92020 American Physical Society simple model for cavity filling in a chambered doctoring system (Fig. 1) and can be connected to nanoimprinting [11]. We note that filling of partially filled cavities could also be an important issue because liquid in the cavities may not be completely emptied in actual gravure processes with continuous operation [10]. Although an array of features is typically used in practice [10,32,33], the model problems considered here provide a rational basis for understanding the interactions between flows in adjacent features. We also note that while the dynamics of contact-line motion can be influenced by three-dimensional (3D) effects [34\u201336], the study of two-dimensional (2D) flows is useful for building physical understanding and is less computationally intensive"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000094_s12283-019-0314-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000094_s12283-019-0314-5-Figure1-1.png",
        "caption": "Fig. 1 Golf club segment definitions and corresponding reference frames. The UpperShaft and LowerShaft reference frames were calibrated with respect to the tracking markers positioned approximately as shown. The ClubFace segment is considered rigidly attached to the LowerShaft segment. The World frame\u2019s X axis points directly at the target and its Z axis is perpendicular to the ground",
        "texts": [
            "5, Vicon Motion Systems, UK) was designed specifically for recording golf swings and utilized eight optical cameras (T40S, Vicon Motion Systems, UK) to record the motion of a club fitted with reflective markers, at a frequency of 723 Hz. The system included two additional synchronized video cameras (Bonita, Vicon Motion Systems, UK) that were used to locate the golf ball before every shot. A total of 20 golfers participated in the experiment. The mean handicap \u00b1 one standard deviation ( ) was 3.6 \u00b1 3.4. For kinematic calculations, the golf club is broken down into three segments: UpperShaft, LowerShaft and ClubFace (see Fig.\u00a01). The UpperShaft and LowerShaft segments have parallel coordinate systems during static calibration, but are considered to move independently during the golf swing. The ClubFace segment is considered rigidly attached to the LowerShaft segment, and so the motion of the LowerShaft segment directly influences the motion of the ClubFace segment. A three-marker cluster was placed on the shaft below the grip to track the movement of the UpperShaft. Four markers were placed on the LowerShaft and ClubFace segments to track the motion of the LowerShaft segment independently of the UpperShaft segment: one marker was placed on the shaft just above the hosel, and the other three were placed on the crown of the clubhead. The approximate locations of tracking markers are shown in Fig.\u00a01. Each rigid segment has a corresponding reference frame. The reference frames were statically calibrated with respect to the tracking markers using numerous calibration fixtures. The ClubFace frame origin is positioned at center-face (CF), with the XC axis normal to the clubface and YC axis parallel with the grooves on the clubface, directed towards the golfer. The UpperShaft and LowerShaft frame origins are located on the shaft axis. During the static calibration, the ZU and ZL axes are coincident with the shaft axis and point toward the butt end of the club, and the XU and XL axes lie in a plane parallel to the XCZC plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002993_s12206-021-0629-6-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002993_s12206-021-0629-6-Figure3-1.png",
        "caption": "Fig. 3. Main parameter schematic of worm drive.",
        "texts": [
            " Furthermore, the parameter calculation methods of worm drive are listed in Table 1 [24], indicating that for this special worm drive, besides a, Z1, i12, mt, all of other parameters can be associated and determined by coefficients k1 and k2, where k1 is the coefficient of base circle of tool rest (also reference circle of worm wheel), and k2 is the coefficient of reference circle of worm. At this point, parameter sensitivity analysis only needs to analyze the influences of k1 and k2. According to previous studies and engineering experiences, the value ranges of k1 and k2 are suggested, given in Table 1 [24]. To clarify the geometrical relationship of main parameters listed in Table 1, the schematic of the worm drive is shown in Fig. 3. The relationships among working condition parameter, geometrical parameter, modeling parameter and coefficients of k1 and k2 are described as Fig. 4. It demonstrates that all of other parameters are associated with coefficients of k1 and k2, thus parametric analysis and optimization study will be conducted around these two parameters. Different values of k1 and k2 will result in various modeling parameters and meshing behavior. Therefore, the effects of k1 and k2 on the meshing performance of the worm drive must be discussed to deal with the parametric analysis of meshing performance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001475_med48518.2020.9183020-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001475_med48518.2020.9183020-Figure1-1.png",
        "caption": "Fig. 1: Two-Dimensional missile-target engagement",
        "texts": [
            " \u2022 On application of the event-based guidance scheme, the stability analysis of the closed-loop system ensures asymptotic stability with Zeno-free behaviour. \u2022 For the proposed guidance scheme, simulations studies on tail-chase and head-on engagement scenarios are carried out to demonstrates its efficiency. In this section a detailed description of the mathematical model of the missile-target engagement is presented, followed by the problem statement. The two dimensional missile-target engagement scenario is represented in Fig. 1. For the point mass system considered, the instantaneous position of the target and missile in the inertial rectangle co-ordinate frame of reference are 978-1-7281-5742-9/20/$31.00 \u00a92020 IEEE 939 Authorized licensed use limited to: Cornell University Library. Downloaded on September 06,2020 at 04:19:48 UTC from IEEE Xplore. Restrictions apply. represented as RT = (RT1, RT2) and RM = (RM1, RM2) respectively. The components along the axis 1 and 2 are represented by the subscripts of \u20321\u2032, \u20322\u2032 respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002444_s43236-021-00220-0-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002444_s43236-021-00220-0-Figure13-1.png",
        "caption": "Fig. 13 Experimental platform",
        "texts": [
            " Figure\u00a012a, b show the torque responses of the motor with and without the UMRP compensation, where both of them are loaded at t = 0.3\u00a0s. As can be seen, the starting torque of the motor is large and the error is larger after stabilization without compensation. The torque response of the motor is more rapid, the torque ripple is smaller, and the rotor suspension performance is better under the improved UMRP compensation. 1 3 Experimental research is carried out using the digital system experimental platform illustrated in Fig.\u00a013. In the experiments, a TMS320F2812 is used as the control chip, a photoelectric encoder is used to detect the motor speed, and an eddy current sensor detects the radial displacement of the rotor. The experimental parameters of the motor are the same as the simulation parameters. Figure\u00a014 shows an experimental block diagram of the control system. Figure\u00a015a shows an experimental diagram of the rotor motion trajectory without compensation. Figure\u00a015b, c show experimental diagrams of the rotor motion trajectory with the UMRP compensation and the improved UMRP compensation, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure5-1.png",
        "caption": "FIGURE 5. Von Misses stress contour",
        "texts": [
            " From the results of static simulation, it can be seen the value of static deflection on the structure of the electric bus frame. The maximum deflection value is 756,69 mm while for the minimum deflection it is 84.06 mm and the average static deflection value is 420.03 mm. After the static deflection value was obtained from the static loading simulation results, then the impact factor values are calculated. This impact factor will be used as a first step in dynamic simulation. Another output of static simulation is the von Misses stress. The von Misses stress result was shown in Figure 5. The maximum stress is 215 MPa and occurred at the joint between roof frame and body frame and also at the joint between body frame and floor frame. The yield stress of material JIS 3445 STKM 13A is 215 MPa, therefore the stress reached the yield stress point, it means the plastic deformation already occurred in the structure. 030153-4 The distance of intrusion of residual space is also the requirement of ECE 66 standard. The value of intrusion must be negative for the safe structure during the rolling accident. Figure 6 shows the value of intrusion distance for the existing material. According to result in the Figure 5 and Figure 6, the frame must be modified to enhance its performance since the structure can be considered as fail. The value of intrusion and also the stress were not fulfilling the requirement of ECE 66 standard. Some part of the frame has been modified by increasing its thickness to reduce the stress and also the value of intrusion. List of modification was tabulated in Table 3. The location of modified part in the frame were depicted on Figure 7. 030153-5 The simulation was done for the modified frame and the result for displacement, stress and also intrusion were depicted in Figure 8, Figure 9 and Figure 10 respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002938_012002-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002938_012002-Figure2-1.png",
        "caption": "Figure 2. (a) The temperature diagram of traditional cooling mold insert, (b) the temperature diagram of conformal cooling mold insert, (c) the cross-section temperature diagram of traditional cooling mold insert, (d) the cross-section temperature diagram of conformal cooling mold insert",
        "texts": [
            " In order to improve the calculation efficiency, a time step of 0.1 s is then selected. The discrete method of simulation time is: 5 s = 0.01 s\u00d720+ 0.1 s\u00d748. After checking the independence of the grid and the time step, the simulated calculation results are obtained. After 5s of cooling, the overall temperature of the conformal cooling mold insert and the traditional MEMAT 2021 Journal of Physics: Conference Series 1939 (2021) 012002 IOP Publishing doi:10.1088/1742-6596/1939/1/012002 cooling mold insert is shown in the Fig. 2. In order to compare the cooling effects of the two cooling methods, a total of 18 points A1-A9 and B1-B9 are selected at the same position on the mold surface, as shown in the Fig. 2 (a) and (b). The temperature of the mold surface point of the two cooling methods is shown in Table 1. The following is a comparative analysis of the cooling effect of the two methods from the aspects of cooling effectiveness, efficiency and uniformity. The conformal cooling channels are directly distributed inside the mold surface to be cooled, which is more effective in cooling the mold surface. According to the data in Figure 2 and Table 1, it can be seen that the mold surface temperature of the conformal cooling is much lower than that of traditional cooling. The average temperature of the conformal cooling mold is 423.88 K, while the average temperature of the traditional cooling mold is 919 K. Because the distribution of traditional cooling channels is limited by manufacturing methods, the distribution of channels is far away from the mold surface that needs to be cooled. It can\u2019t cool down the mold surface in a targeted manner",
            "12 In order to compare the cooling efficiency of the conformal cooling channels and the traditional cooling channels on the mold surface, the cooling efficiency formula is defined as: 0 0 = 100% T T T    (1) \u03b7 means cooling efficiency, T0 represents the initial temperature, and T means the temperature after cooling. The initial temperature is 923.15 K. The effective cooling efficiency of the conformal cooling channels for the mold surface to be cooled can be obtained from formula (1) to be 54.08%. The effective cooling efficiency of the traditional method is only 0.45%. Since the distribution of the conformal cooling channels is closer to the surface to be cooled, the effective cooling efficiency of the conformal cooling far exceeds the traditional cooling method. It can be seen from Fig. 2 (a) that the surface cooling of the mold with traditional cooling channels is extremely uneven. The part close to the cooling water channels has been cooled to about 600 K, while the part far away from the cooling water channels still maintains a higher temperature. Compared with the mold temperature diagram with conformal cooling channel shown in Fig. 2 (b), the cooling effect of the mold surface that needs to be cooled is better, all reaching below 600 K. The temperature gradient diagram of the central section of the water channel is shown in the Fig. 2 (c) and (d). The temperature of the conformal cooling section is mostly distributed in the range of 500-700 K, and the temperature of some areas exceeds 700 K. The temperature of the traditional cooling section is divided into two parts: the high temperature area and the cooling area. The temperature in the area to be cooled reaches 900 K. The traditional cooling mold have large temperature gradients and poor cooling uniformity. Based on the design analysis of the above-mentioned conformal cooling channels, the mold insert was manufactured by XDM 250 from Suzhou XDM 3D Printing Technology Co"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002879_s40032-021-00710-x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002879_s40032-021-00710-x-Figure1-1.png",
        "caption": "Fig. 1 Kinematic diagram of four-links One-DOF EGTs [22]",
        "texts": [
            " Although, various applications of graph isomorphism discussed by several authors as Computer-aided design (CAD), Graphical modeling, Image restoration, Neural networks, Protein structure, Chemical bonding technique, and Social network. The current work establishes checking the isomorphism in EGTs by the method of Wiener number and chemical structure technique [39\u201342]. It is an analytical and less time-consuming technique derived from the adjacency matrix. However, the kinematic structure of a mechanism contains the essential information about which link is connected to which other links by what type of joint. The kinematic structure of a mechanism can be represented in several different ways. Figure 1 shows components of basic epicyclic gear trains such as 1-arm, 2-sun gear, 3-planet gear, and 4-ring gear. Figure 2 shows a functional schematic of four-links One-DOF EGTs. The graphical representation of epicyclic gear trains can be denoted by using two lines, one is a continuous line, and the other is a dashed line. The continuous lines are used for lower pairs as a revolute pair, and dashed lines are used for higher pairs as a gearing pair (Fig. 3). Adjacency matrix, Connectivity matrix, Incidence matrix, and Distance matrix are generally used to represent a conventional graph (or sometimes also called a kinematic graph or skeleton graph) of planar kinematic chains and epicyclic gear trains [22, 23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001374_aim43001.2020.9158987-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001374_aim43001.2020.9158987-Figure4-1.png",
        "caption": "Fig. 4. Planar view of the RRRP manipulator mounted on the platform.",
        "texts": [
            " Then, one could for example prioritize the arm movement instead of base movement, by allocating only a limited portion of the total flow budget for the drive motors. In this study, we only consider the platform flow bounds analytically. For considering arm actuation limitations, we take a different approach, relying on purely tracking error based trajectory adaptation similar to [3]. 627 Authorized licensed use limited to: Cornell University Library. Downloaded on August 29,2020 at 02:41:59 UTC from IEEE Xplore. Restrictions apply. In this simulation study, the manipulator mounted on the 4WS platform is a RRRP crane (see Fig. 4), a topology commonly found in forestry and loader cranes; dimensions of the boom can be found in [11]. The pose of the endeffector in world coordinates is expressed as p = f(q), where the configuration space of the mobile manipulator is q = [ qb qa ] = [ \u03c6 x y q1 q2 q3 d4 ] . To acquire differential kinematics, the base and manipulator Jacobians can be combined into a composite nonholonomic mobile manipulator Jacobian, expressed as p\u0307 = JbGu+ Jaq\u0307a (16) where Ja = \u2202f(q) \u2202qa is the analytical Jacobian of the arm, and G(\u03c6) is a mapping to feasible motions that explicitly models the non-holonomity of the base [12]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure1-1.png",
        "caption": "Figure 1. Location (a) and typical design (b) longitudinal jointing of MC-21 aircraft",
        "texts": [
            " One of the most common types of joints is riveting, bolt riveting and bolt joints, which make up to 90% of the total number of joints of parts [1, 2, 3], as well as in all aircraft. When assembling the MS-21 aircraft, impact riveting technology is prohibited, and in the absence of approaches for the assembly of parts, bolted and riveted bolt connections are used. In particular, most commonly used when connecting the panels along the longitudinal joints in the process of assembling the compartments (Fig. 1) and when connecting the panels along the orbital joints in the process of assembling the fuselage (Fig. 2), bolt-ribbing and, less often, bolted joints. ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 The design of the stud bolt and bolt connections varies depending on the type of rod and bolt head, the size of the head, the material used for the fixing element, and the method of installation in the hole. Examples of rivet and bolt connections are shown in Figure 3 [7]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001376_cpe-powereng48600.2020.9161384-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001376_cpe-powereng48600.2020.9161384-Figure2-1.png",
        "caption": "Fig. 2: Schematic of dual excited synchronous generator",
        "texts": [
            " The extracted mechanical power of the wind turbine could be expressed as a function of wind speed as following [10]: 32 ),( 2 1 wp VCR t Where Pt is the extracted mechanical power from the wind (W), \u03c1 is the density of air (kg/m3), R is the radius pf the turbine blade (m), Vw is the wind speed (m/s) and Cp is the power conversion coefficient which is a function of the blade pitch angle \u03b2 (deg.) and the tip-speed ratio (TSR). The TSR could be defined in terms of wind speed and the rotational angular rotor speed of the turbine. w V R t    \u00a0 \u00a0 \u00a0 \u00a0 \u00a0 (2) Where \u03c9t is the rotational angular speed of the turbine which is associated with the rotational angular speed of the generator \u03c9m and the gearbox ratio (GR) by GRm t  \u00a0\u00a0 \u00a0 \u00a0 \u00a0 (3) The maximum value of Cp (Cp_max) is achieved for \u03b2=0\u00b0 and \u03bbopt. The maximum turbine power is found at a point of \u03bbopt and Cp_max. Fig.2 shows the schematic diagram of DESG having two identical field windings, one of them is located on the d-axis and the other on the q-axis. The Mathematical model of DESG in rotor reference frame can be written in space phasor form under the assumptions of neglecting core saturation, neglecting iron and mechanical losses, and assuming the MMF in the air gap is pure sinusoidal as follows: 353 Authorized licensed use limited to: Auckland University of Technology. Downloaded on August 12,2020 at 12:45:56 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002848_5.0043911-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002848_5.0043911-Figure3-1.png",
        "caption": "FIG. 3. Definition of the quantities used to parameterize the twisted mesoscopic particle structures. The twist axis coincides with the z\u0302 direction, along which also the external magnetic field H is applied. All particles take a distance \u03c1 from the twist axis. From one particle i to the next particle i + 1 above it on the same twisted chain-like particle aggregate, the vertical distance along z\u0302 is denoted as h, while the actual distance between the two particles is referred to as d. The twist angle between the radial position vectors of any such two vertically neighboring particles is called \u03b3.",
        "texts": [
            " All particles are organized in chain-like aggregates, and, for simplicity, they are assumed to take the same distance \u03c1 from the cylinder axis (see Fig. 2). In the initial nonmagnetized state of the system, the particle chains are homogeneously twisted around the cylinder axis. The twist angle \u03b3 quantifies how the radial position vector of the particles is rotated around the cylinder axis from one particle to the next upper one. In the vertical direction, all particles are separated by the same vertical distance h (see Fig. 3). It follows that the distance d between two adjacent particles is given by d = \u221a h2 + 4\u03c12 sin2\u03b3 2 . (5) There are two types of contributions to the magnetization of each particle. First, the external magnetic field H leads to a direct magnetization and a resulting magnetic dipole moment m0 = 4\u03c0 3 a3 3\u03c7 \u03c7 + 3 H. (6) Here, a denotes the radius of the spherical inclusions. The additional factor of 3/(\u03c7 + 3) is associated with the demagnetization effect in a spherical particle.28,57 The second type of contribution to the magnetic moment arises from the mutual magnetization between the magnetized particles",
            " (2), the strain tensor in cylindrical coordinates reads \u03b5 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d \u2212 1 2 A 0 0 0 \u2212 1 2 A 1 2 \u03c4r 0 1 2 \u03c4r A \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 , (27) with the first, second, and third rows or columns being associated with the components along r\u0302, \u03c6\u0302, and z\u0302, respectively. Similarly, we obtain for the magnetization M =M\u03c6\u03c6\u0302 +Mz z\u0302. (28) As before, we do not take into account any anisotropy of the purely elastic contributions to the macroscopic free-energy density. For small enough volume fractions and a significant separation between the magnetizable particles as in our mesoscopic model picture (see Fig. 3), this approximation appears justified. Thus, using \u03b5 from Eq. (27), we obtain the elastic part of the free-energy density of the form Fel = \u03bc \u03b5 : \u03b5 = 3 2 \u03bcA2 + 1 2 \u03bcr2\u03c42. (29) From here, we find the same result for the purely elastic part Fel of the macroscopic free energy as in Eq. (3). Next, we turn to the purely magnetic part of the macroscopic free-energy density, which reads FM = 1 2 \u03b1 : MM \u2212 \u03bc0 M \u22c5H. (30) The structural anisotropy enters this expression via \u03b1 = \u03b1\u2225n\u0302n\u0302 + \u03b1 (I \u2212 n\u0302n\u0302), (31) with the two scalar material coefficients \u03b1\u2225 and \u03b1 as well as the identity matrix I",
            " 4 shows the dependence of the macroscopic system parameters \u03b1\u2225 and \u03b1 [see Eqs. (47) and (48)] on q0. As expected, the corresponding curves are symmetric with respect to the axis q0 = 0. From Eq. (34), it is obvious that the parameters \u03b1\u2225 and \u03b1 by themselves play the roles of inverse magnetic susceptibilities if we confine ourselves to the linear regime. It is then illustrative from the mesoscopic picture that for small values of \u2223q0\u2223, the magnitude of \u03b1\u2225 is smallest (i.e., the corresponding magnetic susceptibility is largest) for q0 = 0. In this case, as Fig. 3 indicates, \u03b30 = 0, and therefore, the particles form vertical chain-like aggregates in the initial state. Their separation distance d is then smallest, which supports their mutual magnetization. Conversely, for small values of \u2223q0\u2223, the magnitude of \u03b1 is largest (i.e., the corresponding magnetic susceptibility is smallest) for q0 = 0. This can likewise be illustratively understood from the mesoscopic picture. When the chain-like aggregates are magnetized perpendicular to their axes, the mutual magnetization between neighboring particles counteracts their net magnetization",
            "60\u201362 The mapping applies as long as electric leakage currents and associated dynamic effects upon electrically polarizing the systems are negligible. The author appreciates stimulating discussions with Dr. G\u00fcnter K. Auernhammer and Professor Stefan Odenbach. Moreover, the author thanks the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) for support of this work through Heisenberg Grant No. ME 3571/4-1. In the following, we further elucidate the changes in the mesoscopic model parameters h, \u03c1, and \u03b3 (see also Fig. 3), as listed by Eqs. (11)\u2013(13). Our expressions are derived to linear order in the elastic deformations as identified by the amplitudes A and \u03c4. For simplicity, we assume that in Cartesian coordinates, the position vector ri of the ith particle in the initial, undeformed state of the system points into the direction x\u0302. Accordingly, ri = \u239b \u239c \u239c \u239c \u239c \u239d \u03c10 0 0 \u239e \u239f \u239f \u239f \u239f \u23a0 . (A1) Consequently, for the (i + 1)th particle, the position vector in the initial undeformed state is given by ri+1 = \u239b \u239c \u239c \u239c \u239c \u239d \u03c10 cos \u03b30 \u03c10 sin \u03b30 h0 \u239e \u239f \u239f \u239f \u239f \u23a0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002216_s12541-020-00457-y-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002216_s12541-020-00457-y-Figure16-1.png",
        "caption": "Fig. 16 FCU test bench",
        "texts": [
            " Because all links are not perfectly stiff, the compliance of each link affects the simulation results of the clamping range. The total combined spring constant of the system was 11 180\u00a0N/mm which is close to the optimal result from dimension synthesis. Since the FCU prototype can only sense clamping force, an FCU test bench was also built to sense the input torque. A 3-D plot of the input torque was drawn with paired panel thickness and clamping force data. The FCU test bench has the same linkage as the FCU itself (Fig.\u00a016). A torque meter was mounted between the handle and the input joint, indicating the data on the upper display. A load cell sensed the clamping force and indicated the data on the lower display. The panel thickness was controlled by the use of different combinations of thickness gauge. The input torque was generated by turning the handle. The 3-D plot of panel thickness vs. input torque vs. clamping force is presented as Fig.\u00a017. Experiment results are plotted as purple dots (left) and simulation results are plotted in various colors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002903_j.mechatronics.2021.102593-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002903_j.mechatronics.2021.102593-Figure1-1.png",
        "caption": "Fig. 1. Design and Mechanical structures of SMP.",
        "texts": [
            " The optical sensing system with three passive joints is developed to measure and control the orientation. A set of three optical sensors is used to correct an IMU error with the KF in order to estimate the cockpit orientation. Furthermore, the SMC incorporating ADC for control operation is implemented to deal with disturbances and compensate for the undesired coupled rotation. In Section 5, The experimental results prove the design and control performance of the SMP for six-DOF motion, including unlimited rotation. Fig. 1 shows the SMP composed of a cockpit sphere for a human on board and a base system for motion control. The base system is designed with four spherical wheels and four linear stages to control the six-DOF motion of the cockpit. It has two redundant actuating systems, but the design improves motion capability and geometrical stability from a larger eccentric margin [18] than the platform with three wheels [17]. The spherical wheel consists of a small sphere, three ball plungers for smooth motion, a passive roller, and an active roller driven by a motor. It is designed with compliant mechanisms to secure four contacts between the cockpit sphere and the four small spheres. Four loadcells are installed on each spherical wheel, and it is located under the spring to measure distributed cockpit weight, as shown in Fig. 1b. The measured data is utilized to estimate and balance the CG of the cockpit sphere. Lastly, each linear stage for translational motion is installed on a ball screw actuator and two linear guides to resist vibrations and decrease loads to the ball screw actuator. The cockpit sphere rotates according to the rotation of small spheres and spherical wheels. The small sphere rotates along with two directions perpendicular to each other. One is free-rolling, enabling the cockpit to move smoothly for both rotation and translation, and a motor drives the other",
            " The failure of the contact eventually incurs degrading control performance. The spherical wheels based on compliant mechanisms are designed to keep four small spheres in contact with the cockpit by compression springs. If the contact occurs in only three points, the spring installed below the ball plunger is compressed by distributed cockpit weight. Then, the small sphere goes down, and a roller still can push up the small sphere by the decompressed spring. Furthermore, induced slip motion between the small sphere and active roller can be minimized, as shown in Fig. 1b. Finally, three small spheres go down by the compliant mechanisms until the cockpit sphere touches the remaining small sphere achieving four contacts. The cockpit can linearly travel along four linear stages installed on the ball screw actuators with a 90\u25e6 separation angle in Fig. 1a. The translational motion is driven from the ball screw actuators with changes of the contact points between the cockpit and the small spheres. Free rolling of the small sphere in the spherical wheel can help the cockpit sphere to move smoothly but consequently causes the cockpit to spin opposite to free-rolling and then undesired coupled rotation between rotation and translation in Fig. 1c. The coupled rotation angle is inversely proportional to the radius ratio between the cockpit and small sphere and should be compensated by counter-rotation from the ADC. The spherical wheels enable the three-DOF rotational and three-DOF translational motions of the cockpit sphere smoothly. The six-DOF motion of the cockpit is analyzed from the kinematics and dynamics where an inertial frame with XIYIZI coordinates is defined at the intersection of four linear stages on the base, as shown in Fig. 1a. Four small spheres on each stage should maintain contact with the cockpit sphere. The necessary condition for the contact is derived from that a circumcenter of a quadrangle P1P2P3P4 is identical to the center of the cockpit in the XIYI plane, which can be expressed in (1). C1 = q1q3 \u2212 q2q4 (1) P0 = 0.5[q1 \u2212 q3 q2 \u2212 q4 2h + 2H1] T (2) where qi (i=1,2,3,4) represents the travel distance of the ith linear stage. R ->R_C and rs are the radii of the cockpit and small spheres, respectively. The position of the center of rotation (CR) P0 can be calculated from the position of the small sphere center Pi and circumcenter equation to satisfy that C1 = 0 in (1)",
            " \u03a9 = NR(q)[I4\u00d74 \u2212 \u039b]q\u03072\u23df\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 \u23de\u23de\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 \u23df rotation with slipping + NT(q)q\u03071\u23df\u0305\u0305\u0305\u0305\u23de\u23de\u0305\u0305\u0305\u0305\u23df coupled rotation (9) where [\u039b]=diag(\u03bb), \u03bb (= [\u03bb1 \u03bb2 \u03bb3 \u03bb4]T) is the slipping ratio satisfying that 0 \u2264 \u03bbi \u2264 1 (i=1,2,3,4). \u03bbi indicates pure rolling (= 0) and slipping (= 1), respectively; rw is the radius of the active roller; q2 = [q5 q6 q7 q8]T, the rotational angle of active rollers and q\u03072 is the angular speed of the rollers. In (9), NR(q) and NT(q) represent Jacobian matrices for rotation and coupled rotation, respectively. The spherical wheels help the cockpit sphere smoothly move but cause undesired coupled rotation \u03a9T due to the translational motion, as shown in Fig 1c, and it is represented that NT(q) = [\u03d2 \u00d7]M(q), where \u03d2 = RC(RC+rs)\u20121[q3\u2012q1 q4\u2012q2 \u20124H1]T. The slip ratio \u03bb in (9) can be estimated from the observer since it is difficult to measure it directly. The state vectors are defined as \u03c71 = [ \u03b4\u03d5\u0307 \u03b4\u03b8\u0307 \u03b4\u03c8\u0307 ] T and \u03c72 = [\u03c71 T \u03bbT]T, where \u03c71 represents velocity error caused by slipping motion. The discrete dynamic model is represented as \u03c71,k+1 = A\u03c71,k+N\u03bbk = C\u03c72,k+1 = CAs\u03c72,k. Then, the observer is designed from (10). \u23a7 \u23aa \u23aa\u23a8 \u23aa \u23aa\u23a9 \u03c7\u0302 2,k+1 = A\u03c7\u0302 2,k + \u03ba3rk+1 rk+1 = [CAs] T[ [CAs][CAs] T]\u2212 1( \u03c7 1,k+1 \u2212 \u03c7\u0302 1,k+1 ) \u03c7\u0302 1,k+1 = CAs \u03c7\u0302 2,k \u03c7 1,k+1 = NR(q)q\u03072 \u2212 \u03a9 + NT(q)q\u03071 (10) where As = [ A N 0 I4\u00d74 ] ,A = I3\u00d73,C = [I3\u00d7303\u00d74],N = NR[diag(q\u03072)] The error for the estimated \u03c72,k is \u03c7\u223c2,k = \u03c7\u0302 2,k \u2212 \u03c7 2,k, and the error dynamics become stable for a positive gain \u03ba3",
            " For the sensor fusion of both the IMU and optical sensors, the cockpit orientation is estimated from a gyroscope. The estimated x2 is corrected from optical sensors for \u03c8 as well as the accelerometer for \u03d5 according to (19). Then, the accelerations ax and ay are measured from the accelerometer. The angular velocity is \u03a9 = [\u03a9X \u03a9Y \u03a9Z]T measured from optical sensors in (19a). As a result, the cockpit orientation can be estimated from both IMU and optical sensors minimizing accumulation error and maximizing the robustness of the optical sensor. The SMP in Fig. 1 is designed with the assumption that PG (CG) is aligned with P0 (CR). However, it is impossible to satisfy due to weight of an operator and equipment. PG should be estimated to calculate G in (13) and apply it into the control system. The unbalanced weight caused by the mismatch between PG and P0 incurs overloads for particular actuators, accelerating slipping motion. The loadcell sensor is implemented to measure transferred mi from the compression spring, as shown in Fig. 3a, where a cover protects the sensor in Fig",
            " This work was partially supported by Development of Multi-degrees of freedom Spherical Motion Platform (2.210065.01) of Institute of Civil-Military Technology Cooperation, the National Research Foundation of Korea(NRF) grant funded by the Korea government (MSIT) (No.2020R1F1A1075857), and Basic Science Research Program through the National Research Foundation of Korea(NRF) funded by the Ministry of Education (No. 2020R1A6A1A03040570) respectively. S.-M. Lee and H. Son Mechatronics 77 (2021) 102593 In Fig. 1a, consider two triangles composed of P1P2P3 and P1P3P4, and circumcenters P0,1 and P0,2 for each triangle can be expressed as follows: P0,i = [ qi \u2212 qi+2 2 (\u2212 1)i+1(qiqi+2 \u2212 q2 2i) 2q2i h ]T , i = 1, 2 (A. 1) It is noted that the full rolling contact between the cockpit and four small spheres can be achieved when P0,1 = P0,2. Then, from (A. 1), the following relation holds 0 = q\u2212 1 4 ( q1q3 \u2212 q2 4 ) \u2212 q\u2212 1 2 ( q2 2 \u2212 q1q3 ) = (q2 + q4)(q1q3 \u2212 q2q4) (A. 2) Since (q2 + q4) \u2215= 0, one has q1q3 = q2q4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000210_ecce.2019.8913105-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000210_ecce.2019.8913105-Figure1-1.png",
        "caption": "Fig. 1: Illustration of the cube sample and the designated X-, Y-, and Z- directions.",
        "texts": [
            " Here, an experimental anisotropy test bed is developed to quantify the B \u2212H performance of the printed iron silicon in three dimensions. Identical B\u2212H characteristics in any direction of excitation is one indicator of true magnetic isotropy. The main objective of the magnetic anisotropy characterization is to obtain the magnetization curves under excitation in the X-, Y-, and Zdirections. This characterization provides the coupled relationship between the magnetic field intensity Hx, Hy , and Hz and the magnetic induction Bx, By , and Bz . A cubic sample with equal cross sectional area on each face, as shown in fig. 1, will allow independent excitation in each direction. The magnetization curve for each direction can be obtained when excitation is applied perpendicular to the cube face of interest. Particularly, the Bx\u2212Hx relationship can be acquired when the excitation is applied to the YZ-plane of the cube. The proposed anisotropy characterization requires a magnetic core made of steel laminations wound with primary windings 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 745 to excite the cube. A general description is shown in fig",
            " There is an airgap on each side between the cubic sample and the magnetic core. As a result, an open magnetic circuit is used for characterization. The primary windings provide excitation to the test specimen. The excitation is controlled to achieve sinusoidal voltage on the secondary windings. Form factor of the secondary voltage, (1), which is the ratio between the RMS value and the average rectified value, should remain within the recommended band of 1.111\u00b11% during the magnetic characterization. FF = V2,rms |V\u03042| (1) For each direction in fig. 1, the corresponding magnetic induction, B, is calculated as in (2). Here, A is the surface area of the face of interest, N2 is the number of turns in the secondary windings wound on the cube, f is the excitation frequency, and |V\u03042| is the average rectified value of the voltage picked up by the secondary windings. B = |V\u03042| 4 \u00b7N2 \u00b7A \u00b7 f (2) The magnetic field intensity, H , of the cubic sample is calculated as in (3), following the magnetic circuit model in fig. 3. Here, Hag is the magnetic field strength in the airgap, Hcore is the magnetic field strength in the magnetic core, lag is the airgap length, lcore is the average length of the flux path within the magnetic core, and lcube is the length of the cube sample"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000913_jfm.2020.272-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000913_jfm.2020.272-Figure1-1.png",
        "caption": "FIGURE 1. A torque swimmer. (a) Schematic diagram of torque swimmers and the problem setting. The torque swimmer consists of a membrane that has a spherical reference shape with radius a, a shear elastic modulus Gs and a bending modulus Eb, enclosing a fluid with viscosity \u00b5 and density \u03c1. A torque surface, expressed by the broken line, is placed \u03b5 above the membrane. For simplicity, the viscosity and density of the external and internal fluids are equal. The straight arrow e in the cell indicates the orientation vector, and the curved arrows on the torque surface indicate torque. (b) Three-dimensional shape of the torque swimmer with Ca = 0.1. The arrow indicates the orientation vector.",
        "texts": [
            " In \u00a7 4, we discuss the biological implications of our results, by comparing them to those of our former experiments using Paramecium (Ishikawa & Hota 2006). Finally, we present our conclusions in \u00a7 5. The basic equations describing our ciliate model are similar to our previous study (Ishikawa et al. 2016), so we explain them only briefly here. 2.1. Deformable torque swimmer The ciliate model consists of a membrane enclosing a Newtonian fluid, and the thrust generated by cilia is modelled as torque distributed above the cell body (cf. figure 1a). The simple model includes swimming and deformation, and exerts membrane tension, so that hydrodynamic interactions can be analysed non-trivially. The reference shape of the membrane is assumed to be a sphere with radius a. We define the unit orientation vector, e, along the line connecting two material points that exist at the posterior and anterior poles of the reference shape. The thickness of the cell membrane is small compared to the cell size and curvature radius, such that the membrane can be modelled as a two-dimensional hyperelastic surface, with surface shear elastic modulus Gs, area dilation modulus Ks and bending modulus Eb",
            " Since the traction force acting on the surrounding fluid and the reaction force transmitted to the cell membrane act in opposite directions with the same magnitude ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 02 0. 27 2 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. B ib ili ot h\u00e8 qu e in te ru ni ve rs ita ir e de S an t\u00e9 , o n 30 A pr 2 02 0 at 0 5: 32 :4 2, 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. (cf. figure 1a), the effective and recovery strokes of the ciliary beat are modelled by torques distributed above the cell body. We assume that the torques are exerted at a distance \u03b5 above the membrane, regardless of the membrane deformation; this is referred to as the torque surface. The length scale of \u03b5 is approximately half the length of the cilia. The strength of torque per unit reference area is assumed to be homogeneous and time invariant. The unsteadiness of the ciliary beat and the inhomogeneity in cilia distribution are ignored"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure1-1.png",
        "caption": "Fig. 1. Cross sections of (a) CP machine (CPM-I), (b) CP machine with Halbach array (CPM-II), (c) CP with slotless stator structure, and (d) CP-Halbach array with slotless stator structure.",
        "texts": [
            " 2) The effect of Halbach PM array with different magnetization direction on reducing saturation of iron poles is systematically investigated. Halbach PM array can lead to superior electromagnetic performance without reducing PM pole arc. It presents an extensive guidance for the design of CP structure. 3) The unequal and multistep staggered rotors with magnetic barriers are proposed. Compared with the regular staggered rotor, the proposed CP PMV machine can not only reduce the unipolar leakage flux in the end region but also achieve better electromagnetic performance. The topologies of the CP PMV machines are shown in Fig. 1 (a) and (b). The CPM-I is the machine with regular CP PM rotor, the CPM-II is the proposed machine with CP-Halbach PM array. CPM-I and CPM-II share the same stator topology. To reduce end winding length and copper loss and to enhance fault-tolerant capability, fractional slot concentrated single-layer windings and fault-tolerant teeth are adopted. Furthermore, higher torque capability can be obtained when the pitch angles of adjacent modulator poles are unequal and the distribution of modulator poles is uneven [24]. The detailed parameters of proposed machine are summarized in Table I. Moreover, to reveal the characteristic of CP structure, the CP and CP-Halbach array with slotless stator structures are depicted in Fig. 1 (c) and (d). The air-gap flux density of CP PM machine without considering slotting effect and leakage flux can be expressed as [23]   1 1 r cpm r m B B g h      (1) where Br is the remanence of PM, g is the air-gap length, hm is the PM thickness, \u03b1 is the pole arc coefficient defined as the ratio of PM pole arc to the pole pitch, \u03bcr is the relative permeance. Furthermore, the air-gap flux density of the salient iron core can be expressed as [23] 1 cpr cpmB B     (2) Therefore, it can be found that when \u03b1>0",
            " 2 2 0 0 2 2 s s s T T cpm s cpr s T s B B d B d T            (3)     2 2 0 2 4 cos cos s s s T T i cpm r s s cpr r s s T s B B iP d B iP d T              (4) where Pr is the PM pole pairs, \u03b8s is the mechanical angle, Ts is the period of flux density distribution, which equals to 2\u03c0/Pr. Then, the Fourier expression of ideal airgap flux density of CP structure can be expressed as        0 2 sin , cos 1 r m cp s r s i m r B h i B t B i P t i h i g                (5) It can be found that the fundamental airgap flux density is related to \u03b1, g, hm and PM material. Secondly, actual flux density distribution can be investigated by finite element analysis (FEA). Hence, a CP rotor with slotless stator structure shown in Fig. 1 (c) is built and analyzed. Fig. 3 (a) is the variation of ideal fundamental airgap flux density calculated by (4). It can be found that the optimal \u03b1 increases as hm increases. Although the flux density increases as hm increases, its increase rate decreases. Fig. 3 (b) is the variation of actual fundamental airgap flux density predicted by FEA. Obviously, the optimal pole arc coefficients under ideal and actual distribution are different. Fig. 3 (c) is the variation of leakage flux coefficient, which is defined as 100%cal FEA leakage cal B B C B    (6) where Bcal is the ideal fundamental flux density value calculated by (4), BFEA is the actual fundamental flux density value predicted by FEA",
            " Moreover, when \u03b4 increases, the magnetic circuit is shortened, resulting in much less radial flux lines in salient iron pole, and the diffluence effect is enhanced. Therefore, the flux density in Region-II will be lower than that in Region-I, which will weaken the saturation of salient iron pole. Thus, if the flux density in the Region-I is kept in the normal range, the saturation of salient core can be eliminated completely. To reveal the influence of rotor parameters on the saturation of salient iron pole, a CP-Halbach PM rotor with slotless stator is built and the flux density of the Region-I is investigated, as shown in Fig. 1 (d). It is worth noting that when the stator is slotless, the flux densities in the airgap and the salient iron pole are the largest, leading to the highest level of magnetic saturation. The influence of the rotor parameters on the flux density in region-I with variable \u03b4 is investigated. It can be observed from Fig. 6 that when \u03b1 is increased, the flux density in Region-I becomes higher. However, hm has less effect on the peak flux density. The main reason is that although the MMF has been enhanced due to thicker PMs, the reluctance of PMs is also increased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002738_s12555-019-1032-2-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002738_s12555-019-1032-2-Figure1-1.png",
        "caption": "Fig. 1. Two-wheeled inverted pendulum system model.",
        "texts": [
            " Underactuated system In underactuated systems, where the control input T \u2208 Rm\u00d71 and the matrix of control coefficient B \u2208 Rn\u00d7m, and n > m, we can obtain the constraint force as follows: Qc(x, x\u0307, t) =(M(x, t)H\u22121(x, t)B)+ \u00d7 [b(x, x\u0307, t)+M(x, t)H\u22121(x, t)Q(x, x\u0307, t) \u2212\u03bb (M(x, t)x\u0307\u2212 c(x, t))], (18) where Q(x, x\u0307, t) =C(x, x\u0307, t)x\u0307+G(x, t)+U(x, x\u0307, t), and \u201c+\u201d stands for the Moore-Penrose generalized inverse. However, in the actual system, the uncertainty is usually unknown. Equations (17) and (18) cannot be applied to real systems. Therefore, we design a more practical control to obtain the constraint force when the uncertainty is unknown (see Section 4). 3. DYNAMIC MODEL 3.1. System description In this section, we describe the two-wheeled inverted pendulum system and derive its dynamic equation. The simplified system model is shown in Fig. 1 and the parameters and variables are defined in Table 1. Suppose that there is no slip between the wheels and the ground. We divide the system into two subsystems. The first one is the yaw subsystem with T2 = Tl \u2212Tr as the control input, which is used to control the angle of yaw \u03d5 . This subsystem has one input and one degree of freedom, which is a fully actuated system. The other system is underactuated and has two degrees of freedom, the angle of the inverted pendulum, \u03b8 , and the displacement of the system, x, with T1 = Tl + Tr as the control input"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001716_ccs49175.2020.9231465-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001716_ccs49175.2020.9231465-Figure2-1.png",
        "caption": "Fig. 2. The overall structure of the custom vision-based tactile sensor.",
        "texts": [
            " In this paper, the design of a custom FingerVision is demonstrated, and a method of specific contact surface recognition is to be introduced with optimization of its feature extraction algorithms. An accuracy test with recognition effect analysis is attached to the end as well. II. CUSTOM FINGERVISION DESIGN The basic scheme of obtaining contact information from a vision-based tactile sensor is to visualize the deformation of the material in the contact area by specific means and record it as a visual information through a camera. The overall structure and dimension of our custom sensor is shown in Fig. 2, like the structure of FingerVision, consists of a specific silicone elastomer, a transparent hard acrylic layer, a camera underneath and a framework, which is made through 3-D printing. As for the camera module, we use the commercial camera, Logitech BRIO Webcam, for its capability of 4K Ultra HD video and automatic zooming control. The camera is fixed on the framework through bolts and nuts, and uses USB connection to transmit images or videos to PC. Although the camera stream of BRIO Webcam can achieve 1080P at 30fps, it has a large length of 100mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002129_cdc42340.2020.9304199-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002129_cdc42340.2020.9304199-Figure2-1.png",
        "caption": "Fig. 2. Illustration of one UAV\u2019s path following error.",
        "texts": [
            " During the formation flight, each UAV\u2019s heading angle aims to coincide with the tangent direction of \u0393 at the formation center point, while maintaining a displacement (dxi , d y i )T expressed in the inertial frame with respect to the formation center. Thus it can be regarded that a new path \u0393i derived by adding the displacement (dxi , d y i )T to the original path \u0393 is assigned to UAV i, as shown in Fig. 1. It is assumed that the path \u0393 and the displacement (dxi , d y i )T is known to UAV i. Thus the UAV can be programmed to follow \u0393i while adjusting its progression rate to coordinate with the other UAVs. Now we formulate the coordinated path following problem. As shown in Fig. 2, pi denotes the closest projection point of UAV i on the path \u0393i, and T(pi) denotes the tangent vector of \u0393i at pi. Let \u03c1i \u2208 R represent the distance from UAV i to pi with a sign: \u03c1i > 0 when UAV i is on the left side of \u0393i, and \u03c1i < 0 when it is on the right side. The symbol \u03c8i denotes the heading angle of UAV i with respect to T(pi). To achieve path following under control input constraints, it is assumed that the absolute value of the curvature at any point of the path \u0393 is less than a constant \u03ba0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002881_s42417-021-00329-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002881_s42417-021-00329-3-Figure1-1.png",
        "caption": "Fig. 1 Model of the gear transmission system. a The test gearbox. b 3D model",
        "texts": [
            " Then the fourth part mainly compares the simulation results with the experimental results of the gear transmission system temperature field, which verifies the correctness and effectiveness of the proposed method. Finally, in the fifth part, the influence of load, speed, and other factors on the temperature response of the system is discussed in detail, and the high-temperature response caused by system resonance can be avoided by adjusting the gear position. In this paper, the two-stage gear transmission system is taken as the study object, as illustrated in Fig.\u00a01. The gearbox consists of six deep groove ball bearings, three shafts, and twostage spur gears. 1 3 The analysis process is shown in Fig.\u00a0 2. Firstly, the dynamic model of gear transmission system is solved by newmark-\u03b2 time integration method, the dynamic load and vibration displacement of each element is extracted. Then, combined with the theory of tribology and heat transfer, the heat flux of each heat source and the convective heat transfer coefficient of each part are determined. On this basis, combined with the finite element method, the temperature field of the transmission system model is simulated and calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000894_j.addma.2020.101208-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000894_j.addma.2020.101208-Figure3-1.png",
        "caption": "Fig. 3: a) Sketch of the setup and b) final scanning strategy.",
        "texts": [
            " 2: X-ray diffraction pattern of the GA V-9Si-5B powder material confirming the presence of the phases Vss, V3Si and V5SiB2. However, the argon GA V-9Si-5B powder have satisfying microstructural homogeneity and show up with a predominant spherical shape and a low tendency for agglomeration (visual observation, Fig. 1c) meeting the requirements for DED to achieve constant powder flow rates during build-up and avoid porosity in the compacts [16]. A sketch of the general setup for the DED experiments is depicted in Fig. 3. The laser-beam source was a 2 kW diode laser. A continuous coaxial powder nozzle was used. The experiments were carried out in an inert gas chamber with an oxygen content below 50 ppm. As carrier gas, for the protective inert gas stream of 15 l/min and global shielding argon 4.6 was used. The substrate material was preheated to approx. 710\u00b0C. The induction coil used is freely movable in built direction \u201cz\u201d to ensure a constant preheating temperature in the respective processing plane. r -p ro of The used scanning strategy was an alternating bidirectional strategy with double exposure without powder feed subsequent to the built of every second layer (Fig. 3b). The purpose was to smoothen the surface by re-melting and thus stabilize the build-up of the subsequent layer. The DED parameter used in this study are listed in Tab. 2. The height of each layer was 0.2 mm. Tab. 2: DED parameters used in this study Parameter Symbol Unit Value build-up Value re-melting Laser beam power PL W 175 120 Laser beam diameter D mm 0.6 Scanning velocity vv mm/s 6.67 Overlap \u0394ys mm 0.3 - Powder mass flow VP cm\u00b3/min 0.56 - DED samples of 4 x5 x 10 mm\u00b3 in size were built as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000111_s10846-019-01102-1-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000111_s10846-019-01102-1-Figure1-1.png",
        "caption": "Fig. 1 Body and inertial coordinate systems for the AUV",
        "texts": [
            " The rest of the paper is organized as follows: Section 2 studies the mathematical model for an AUV which represents the equations of motion and the calculation of hydrodynamic stability derivatives coefficients. The design of control system for the AUV is presented in Section 3 and finally the simulation results are presented and discussed in Section 4. The equations of motion for an AUVare usually determined in body coordinate system or relative to a fixed or inertial coordinate system which are depicted in Fig. 1. As Fig. 1 shows, the notations x, y and z are the translational positions of the vehicle relative to the inertial coordinate system and \u03c6, \u03b8 and \u03c8 are the rotational positions of the vehicle relative to the inertial coordinate system. Also, u, v and w are the linear surge, sway and heave components of velocity in the body coordinate system respectively and p, q and r represent the components of angular velocity. The resultant forces exerted to the vehicle in the body coordinate system are represented by X, Y and Z and the resultant moments are expressed as: K, M and N in the body fixed coordinate system",
            " 2, \u03c8c, \u03b8c and \u03c6c indicate the input commands of yaw, pitch and roll respectively. The error values of yaw, pitch and roll which are shown respectively as e1, e2 and e3 are inputs of proportional-integral-derivative (PID) controllers and outputs of controllers are the rudder, elevator and aileron commands which are represented as \u03b4rc, \u03b4ec and \u03b4ac respectively. Commands of four actuators (\u03b41c, \u03b42c, \u03b43c and \u03b44c) are then calculated in the compensator as follows: \u03b41c \u00bc \u03b4ac \u00fe \u03b4ec \u00fe \u03b4rc \u00f020\u00de \u03b42c \u00bc \u03b4ec\u2212\u03b4rc \u00f021\u00de \u03b43c \u00bc \u03b4ac\u2212\u03b4ec\u2212\u03b4rc \u00f022\u00de \u03b44c \u00bc \u2212\u03b4ec \u00fe \u03b4rc \u00f023\u00de As Fig. 1 shows, four control surfaces have a cross orientation with an angle of 45\u00b0 relative to the horizon. The number of each control surface form 1 to 4 is added in a clockwise manner (North-East (1), South-East (2), South-West (3) and North-West (4)) for an observer from the behind of the vehicle. Actuators transfer the input commands to the deflection of four control surfaces (\u03b41, \u03b42, \u03b43 and \u03b44) which are inputs of the system dynamics and finally the yaw, pitch and roll angles of vehicle are the output of system dynamics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002509_14644193211003777-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002509_14644193211003777-Figure2-1.png",
        "caption": "Figure 2. Schematic diagrams of rotor-bearing-housing system.",
        "texts": [
            " The dynamic modeling of rotor-bearing-housing system is presented in section 2. Then, the results of the simu- lation are verified by using the experimental data in section 3. The effect of rotor mass distribution on the vibration response of the bearing is discussed in section 4. Finally, some conclusions are drawn. Dynamic modeling of rotor-bearing-housing system The structure of the electrical motor is shown in Figure 1 of which components include rotor, bear- ings, and motor housing. The system can be simplified as Figure 2(a). The rotor is modeled as a flexible rotor. The stiffness of the shaft is obtained by regarding it as a beam element. The bolted connection between the actual housing and the base plate is simplified by using four sets of linear springs and damp- ers, two sets acting in the X direction and Y direction, respectively. The bearings locate at the fan end and the drive end separately. Each rolling element modeled as a mass element is attached to the inner ring and outer ring via a spring and damper, which is shown in Figure 2(b). In order to simplify the rotor-bearing system, the following assumption are made: 1. The bearing\u2019s inner ring is completely connected with the shaft and the bearing\u2019s outer ring is completely connected with the housing; 2. The model is linearized contact, no contact friction and isothermal; 3. There is no slippage between the rolling elements and the raceway; 4. The rolling elements experience only elastic Hertzian contact; 5. The Hertzian contact forces between the raceway and rolling elements may occur only during compression; 6",
            " Therefore, it is necessary to add a fault model with local defects on the raceway to the dynamic model, and to add a timevarying and position-dependent raceway diameter into the dynamic model, which can be expressed as equations (15) and (16) d1i;i\u00bc1::9 \u00bc rd \u00fe rd;ball qi; qi < rd \u00fe rd;ball 0 qi > rd \u00fe rd;ball d1i;i\u00bc10::17 \u00bc rf \u00fe rf;ball qi; qi < rf \u00fe rf;ball 0 qi > rf \u00fe rf;ball (15) d2i;i\u00bc1::9 \u00bc xi \u00f0Rd rd;ball\u00de; xi < Rd rd;ball 0 xi > Rd rd;ball d2i;i\u00bc10::17 \u00bc xi \u00f0Rf rf;ball\u00de; xi < Rf rf;ball 0 xi > Rf rf;ball (16) where xi is the position of the i-th bearing element relative to the outer raceway, which can be expressed as equations (17) and (18) xi \u00bc f xds xdh \u00fe qicoshi\u00f0 \u00de2 \u00fe \u00f0yds ydh \u00fe qisinhi\u00de2g 1 2 (17) xi \u00bc f xfs xfh \u00fe qicoshi\u00f0 \u00de2 \u00fe \u00f0yfs yfh \u00fe qisinhi\u00de2g 1 2 (18) As shown in Figure 2(b), the width, depth, and angle of the local defect on the outer ring are Wd, Cd, and hd. When a rolling element passes the bearing defect at any given time step, the outer raceway diameter of the rolling element will be adjusted accordingly to reflect the effective increase in the raceway diameter due to the defect located on the outer raceway. Therefore, it is necessary to determine the angular position of the rolling element at any given time step. hi \u00bc 2p\u00f0i 1\u00de Z \u00fe xct\u00fe h0 (19) where Z is the number of rolling elements, i is the serial number of the rolling element, h0 is the initial angle of a rolling element with the serial number 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000652_icces45898.2019.9002393-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000652_icces45898.2019.9002393-Figure1-1.png",
        "caption": "Fig. 1. 4 wheel steering dynamic model",
        "texts": [
            " \u2022 Validation of proposed controller in Carsim software. Organization of this paper is as follow, Section I gives brief introduction of latest work. Section II derives the 4WS nonlinear vehicle model along with desired model of lateral dynamics. An IDC based sliding mode controller is implemented in Section III. Section IV shows the effectiveness of proposed controller and Section V convey robustness through Carsim validation. Section VI is conclusion. A double track model of 4 wheel steering vehicle is seen in Fig.1 . [11], [12] The 2 DOF vehicle model captures the lateral dynamics obtained as mVx(\u03b2\u0307 + r) = Fy1cos\u03b4f + Fy2cos\u03b4f + Fy3cos\u03b4r + Fy4cos\u03b4r +Fx1sin\u03b4f + Fx2sin\u03b4f + Fx3sin\u03b4r + Fx4sin\u03b4r (1) Izz r\u0307 = a(Fy1cos\u03b4f + Fy2cos\u03b4f + Fx1sin\u03b4f + Fx2sin\u03b4f ) \u2212b(Fy3cos\u03b4r + Fy4cos\u03b4r + Fx3sin\u03b4r + Fx4sin\u03b4r) + tw 2 (Fy1sin\u03b4f \u2212 Fy2sin\u03b4f + Fy3sin\u03b4r \u2212 Fy4sin\u03b4r +Fx2cos\u03b4f \u2212 Fx1cos\u03b4f + Fx4cos\u03b4r \u2212 Fx3cos\u03b4r) (2) where \u03b4f and \u03b4r are front and rear wheel angle, ay = vx(\u03b2\u0307+ r) is the lateral acceleration at C.G of vehicle. Fxi and Fyi are longitudinal and lateral forces simultaneously produced by front and rear tires"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001041_iet-epa.2020.0237-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001041_iet-epa.2020.0237-Figure3-1.png",
        "caption": "Fig. 3 Stator core radial dimensions",
        "texts": [
            " Downloaded on October 19,2020 at 07:29:02 UTC from IEEE Xplore. Restrictions apply. in the thermal network in Fig. 2 are calculated and summarised hereafter. Interface gap conduction resistance (R1): The interface gap occurs during the assembling process of packing the stator to the housing, and it is evaluated as [17] R1 = lig kair\u03c0DosLs , (4) where lig is the interface gap between the frame and stator laminations (m), which is selected according to [22] . Stator yoke conduction resistances (R2 and R3): According to Fig. 3, the stator yoke is divided into two sections, and the value of the radial conduction resistance of each part is evaluated by using (2) as follows [17]: R2 = ln ros/rm 2\u03c0kirLs , (5) R3 = ln rm/riy 2\u03c0kirLs , (6) where rm is the mean value of the stator yoke radius (m), roy is the outer radius of the stator, and riy is the inner radius of the stator yoke (m). Stator teeth conduction resistance (R4): To evaluate the equivalent thermal resistance of the stator teeth, the geometry of this motor portion is assumed as a cylinder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000462_s10556-020-00699-7-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000462_s10556-020-00699-7-Figure4-1.png",
        "caption": "Fig. 4. TRIZ damper bearings: (a) \u2014 Schematic diagram of PD support bearing with three pads; (b) \u2014 PDU combined journal and axial bearing in the course of assembly.",
        "texts": [
            " Through the use of oil scrapers the temperature of a loaded bearing support pad could be decreased, though this did not solve the problems associated with subsidence of the rotor in the course of turbine operation. Subsidence of the rotor shaft occurs as a consequence of collapse of the support surface of the pad, which rests on the body of the bearing (linear contact on the back edge of the pad). To solve this problem the TRIZ firm developed the PD and PDU damper bearings each provided with three pads (Fig. 4, a) in order to replace the standard support bearings with mechanical support. The support part of the three-pad bearing developed by TRIZ was calculated (Table 4). The carrying capacity of the support bearings developed by TRIZ (cf. Table 4) is higher than the carrying capacity of standard bearings with five pads (cf. Table 1), a circumstance that may be explained by the greater area of the support surface of the TRIZ bearings. Thus, in the PD support bearing produced by TRIZ the pads occupy 90% of the surrounding space whereas in a bearing with five blocks only 75%",
            " It is of no little importance that with the use of the proposed bearing it is possible to reduce losses of power due to friction (by 22%) and the maximal temperature of pads (by 13%). A significant advantage of the damper bearings produced by TRIZ is the possibility of self-adjustment of the bearing pads in both the radial and transverse directions. This ability assures guaranteed absence of mechanical contact between the pads, low losses of power, and lengthy operating life without replacement of the pads. Since 2002, PD120 bearings produced by TRIZ have been installed on 29 103JT turbines, replacing standard support bearings (cf. Fig. 4, b). The carrying capacity of the standard combined journal and axial bearing of turbine 103JT at the hightemperature end of the rotor does not conform to the operating conditions. Installation (in the overhaul period of the turbine) of oil scrapers is a quick and effective technical solution for reducing the temperature in the loaded pad of a standard bearing with five pads, though it does not yield a solution to the problem of subsidence of the shaft when a standard bearing is used. Through replacement of standard bearings by the damper bearings produced by TRIZ it becomes possible to increase the carrying capacity of the bearings, reduce friction-induced losses of power by 22%, decrease the maximal temperature of the pads by 13%, protect the surface of the pads against electro-erosion damage, successfully solve the problem of subsidence of the rotor, and increase the operating life of the bearing without replacing the pads"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001507_012020-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001507_012020-Figure2-1.png",
        "caption": "Figure 2. Injection molded magnet carrier a) Soft magnetic composite (inner) and Polyamide (outer) b) Magnets mounted on single rotor disc [3]",
        "texts": [
            " It is seen that the weight of stator is 41%, housing 39% and rotor 20%. There are two types of masses: stator and housing do not rotate but only translate and rotor rotates and translates. Rotational inertia reduction is aimed for saving weight. Better dynamic behavior can be obtained by lowering the rotational inertia. While accelerating or decelerating, higher efficiency of the motor is obtained by lowering the rotational inertia [2]. Injection molded magnetic carrier and magnets mounted on a single rotor disc are shown in figure 2. This structure combines two functional requirements: torque transmission and magnetically active part. GFRP (Glass fiber reinforced plastics) is used for transmitting torque and SMC is used for the flow of magnetic field. The length of the magnet carrier was made equal to the length of the magnet. Multiple magnet carriers are assembled to accommodate required number of magnets. Splines are used for easier assembly on the rotor shaft and for a positive torque transmission between the rotor disc and the shaft [3]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002604_j.apacoust.2021.108063-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002604_j.apacoust.2021.108063-Figure2-1.png",
        "caption": "Fig. 2. Determining the modulus of rubber mater",
        "texts": [
            " In reference [16], the parameters of the elastoplastic model were identified by reducing the gap between the tested and simulated moduli of rubber. Similarly, in the current study, the elastoplastic parameters were calculated by narrowing the gap between the dynamic mechanical analysis (DMA) experimental modulus and simulated modulus. Rubber samples of dimensions 17.5 mm, 8.9 mm and 2.5 mm in length, width and height respectively, were cut. In the DMA test, each sample was clamped at one end and imposed with a sinusoidal displacement excitation on the other end using Q800 V7.1 Build 116, as shown in Fig. 2(a). It was finished under a constant temperature (25 C), with the excitation frequency increased from 1 Hz to 50 Hz, at 2 Hz step size. The required signal was gathered by the force and displacement sensors and recorded by a computer. Under the assistant of sample geometries, the required stress and strain could be obtained. Then the modulus was calculated. The data was processed by the Universal Analysis 2000. A series of dynamic simulations were performed with the conditions identical to the tested conditions, as depicted in Fig. 2(b). Fig. 3 reveals a comparison of the experimental complex modulus, the simulated complex modulus with and without considering the elastoplastic model. The simulated complex modulus considering the elastoplastic model agrees well with the experimental results under both 1 mm and 0.3 mm excitation amplitudes. However, without considering the elastoplastic model, the complex modulus is insensitive to the excitation amplitude. The above results demonstrate the validity of the elastoplastic model. The identified parameters are listed in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001102_j.matpr.2020.05.310-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001102_j.matpr.2020.05.310-Figure5-1.png",
        "caption": "Fig. 5. Stress developed in the wheel housing.",
        "texts": [
            " The mesh is then opened in the software for the analysis in which initially the load is applied to check for the stresses. The load applied gets enforced in the housing which is portrayed in the Fig. 3. On applying the stresses in the housing the deflection is noted in the output as 0.42 mm for the applied load as shown in the Fig. 4. The stress acting on the housing of the wheel while applying load is 39.008 N/mm2 as al., FEA based approach on replacing the metal cast wheel into thermoset 20.05.310 shown in the Fig. 5. The values determined were found to be within the safety limits since the maximum permissible value was 118 N/ mm2 mentioning consideration in usage of the made wheel housing. As an outset, considering various calculations and design development for the product and mould design for SMC, wheel housing was completed with actual values as guidance. Based on the needs of the automotive industry the product can be perfected considering the needs. Based on the implementation, the prototype can be made for experimental analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003415_9781119352181.ch4-Figure4.26-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003415_9781119352181.ch4-Figure4.26-1.png",
        "caption": "Figure 4.26 Detailed slot and end-winding configurations for 250 HP machine.",
        "texts": [
            " Stator conductors: Bare 0.129 \u00d7 0.204\u2032\u2032 Insulated 0.146 \u00d7 0.220\u2032\u2032 Stator slot insulation: Width 0.145\u2032\u2032 Depth 0.240\u2032\u2032 Bars in parallel per turn: 1 Distance between coil sides: 1/16\u2032\u2032 Winding pitch: 12/15 = 0.8 Rotor conductors: 3/8 \u00d7 9/16\u2032\u2032 bar Rotor bar extension: 1 5/8\u2032\u2032 Stator coil extension (straight portion): 1 1/4\u2032\u2032 Rotor skew: 1 stator slot pitch Using the above data, the arrangement of conductors in a typical stator and rotor slot can be surmised. Details of stator and rotor slots after winding are shown in Figure 4.26. Note that the stator slot has three types of insulation: (1) insulation between conductors, (2) insulation between coil sides, and (3) insulation to ground. The stator is usually heavily insulated since the potential to ground is generally high (2400 V in this case). The rotor is much less insulated since there is effectively only one turn per slot so that the potential to ground is small. For smaller machines, the insulation material used to isolate the rotor bars from the rotor laminations may be simply an oxide build up intentionally given to either the rotor bars or slots during a heat treatment process. In general, the total rotor copper cross section is selected to be in the range of 60\u201380% of the stator copper cross section for a squirrel-cage machine and 85\u201390% for a wound rotor machine. The limiting factor is heating of the bars which limits the current density during starting to values typically on the order of 45,000 A/in.2 (70 A/mm2). The corresponding values for the stator are in the range of 30\u201335,000 A/in.2 (46\u201354 A/mm2), increasing with machine size. Figure 4.26 also shows details of the stator and rotor end-winding regions. Stator Slot Leakage Inductance From Figure 4.26, note that the clearance from the insulated conductors to the sides of the slot is (0.374 \u2212 0.220)\u22152 = 0.077\u2032\u2032 It will be assumed that the same tolerance is allowed at the bottom of the slot. The distance d6 as defined by Figure 4.8 is equal to this distance plus one insulation thickness for the insulated stator conductor d6 = 0.077 + (0.146 \u2212 0.129)\u22152 = 0.0855\u2032\u2032 The distance d5 measured from the start of the conductor of the bottommost conductor to the end of the conductor in the topmost conductor is d5 = 5(0",
            "36), the slot leakage per bar is Lb = n2 s ler psl = (1)2 ( 8.68 39.37 ) 1.461 \u00d7 10\u22126 = 0.32 \u03bcH\u2215bar Note that the quantity ler takes into account the slight additional length of the rotor bar due to skew, that is, ler = les cos ( 2\u03c0 S1 ) = 8.67 cos ( 2\u03c0 120 ) = 8.68\u2032\u2032 Rotor End-Winding Leakage Inductance per Ring Segment The end winding leakage inductance per segment of the end ring is obtained from Eq. (4.134) as Le = \u03bc0 Db 2S2 { 1 2 [ 1 + 1 6 ( dbetbe D2 b )] ln [ 8 D2 b dbetbe ] \u2212 0.8434+ 0.2041 ( dbetbe D2 b )} From Figure 4.26 the rotor diameter measured at the middle of the end ring is Db = [Dor \u2212 2dsr \u2212 dbe] = [24 \u2212 2(0.67) \u2212 0.75] = 21.91\u2032\u2032 = 55.65 cm Le = (4\u03c010\u22129) ( 55.65 2 \u00d7 97 )((1 2 )( 1 + 1 6 ( 1 \u00d7 0.75 21.912 )) ln ( 8 \u00d7 21.912 1 \u00d7 0.75 ) \u2212 0.8434 + 0.2041 1 \u00d7 0.75 21.912 ) = 0.01251 \u03bcH/end ring segment The next task in the calculation of machine parameters is to relate the inductance per bar just calculated to an expression which involves the overall rotor leakage inductance per phase. This is the topic of Section 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002807_s40430-021-03005-5-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002807_s40430-021-03005-5-Figure4-1.png",
        "caption": "Fig. 4 Spraying pattern of each UAV",
        "texts": [
            " \ufffd Ix\u2212Iy Iz \ufffd \u239e \u239f\u239f\u239f\u239f\u23a0 , Fd = \u239b\u239c\u239c\u239d Fdx Fdy Fdz \u239e \u239f\u239f\u23a0 , \ud835\udf0fd = \u239b\u239c\u239c\u239d \ud835\udf0fd\ud835\udf19 \ud835\udf0fd\ud835\udf03 \ud835\udf0fd\ud835\udf13 \u239e\u239f\u239f\u23a0 (2)I( , , ) = \u239b \u239c\u239c\u239c\u239d 1 Ix 0 0 0 1 Iy 0 0 0 1 Iz \u239e\u239f\u239f\u239f\u23a0 (3) g ( q, pi ) = g ( q \u2212 pi ) = i(t) exp ( \u2212 ( qx \u2212 pxi )2 + ( qy \u2212 pyi )2 2 s ) 1 3 where ri = \u2016q \u2212 pi\u2016 , s is the spread of spraying pattern and i is a positive adjustable value representing the effectiveness of spraying operation for the ith agent. Since the spraying effectiveness will attenuate to zero as \u2016q \u2212 pi\u2016 \u2192 \u221e , one can define the effective region of spraying by \u03a9g (Fig.\u00a04). The parameter i is defined to adjust the spraying effectiveness of each agent, which would be detailed in Sect.\u00a02.3. Due to the limited performance area of the agents, the mission area is first partitioned into subregions assigned to the agents using Voronoi partitioning. Assume that the movement of each robot is confined in Q and that V = { V1,\u2026 ,VN } is a generalized Voronoi tessellation of Q such thatI ( Vi ) \u2229 I ( Vj ) = \ufffd . I(\u22c5) denotes the interior space of each Vi and \u22c3N i=1 Vi = Q . Therefore, it is supposed that each agent is only responsible for covering its domain Vi [22]",
            " This approach is based on the assumption that the closed-loop attitude dynamics converge faster than the closed-loop translational dynamics [41, 42]. Desired references p\u2217 i = ( x\u2217 i , y\u2217 i ) ,z\u2217 i and \u2217 i are generated by a trajectory generator in the outer loop of the proposed framework in Fig.\u00a05, whereas references \u2217 i and \u2217 i are generated by using the dynamic equation of x\u0308i and y\u0308i in Eq.\u00a0(2). We suggest the reader refer to [36] for more details. The control strategy for the UAVs is implemented in a distributed form, as depicted in Fig.\u00a05. The Voronoi-based trajectory generator in the outer loop of Fig.\u00a04. determines the new positions for the UAVs at each time step. To move to this new position, each UAV will use the ITST-SM control in the inner loop. Based on this framework, the closed-loop system converges to a centroidal Voronoi tessellation, and then, the dispersants would be sprayed on the oil spill. Therefore, we (20)\ud835\udf09\ud835\udefc 1 + 2\ud835\udf1a\ud835\udf14?\u0307?\ud835\udefc 1 + \ud835\udf142 ( \ud835\udf09\ud835\udefc 1 \u2212 TVi ) = 0 consider the UAV as a rigid body. Moreover, note that since each robot moves toward its Voronoi center, and no one could leave its Voronoi cell, there is no possibility of collision"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002025_nir50484.2020.9290199-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002025_nir50484.2020.9290199-Figure1-1.png",
        "caption": "Fig. 1: The structure of the quadrotor",
        "texts": [
            " This can be seen from the works [15], [16], which show how complex the dynamics of the quadrotor is, given the interaction between all its nodes. This paper is organized as follows. The addressed problem is set up in Section 2. The considered plant model is parameterized in Section 3. Estimation methods, which are advanced Kalman filter and the DREM, are provided in Section 4. Their application to the parameterized model of the quadrotor motion is illustrated in Section 5. Finally, the main finding are summarized in Conclusions. A quadrotor has a symmetrical structure as it is shown in Fig. 1. Four actuators located on the rays along the axes ObXb and ObZb at a distance l from the center of mass. The inertia matrix has the following form: I = Ixx 0 0 0 Izz 0 0 0 Iyy  , (2) where Ixx, Iyy, Izz are moments of inertia during rotation along the corresponding axes, Ixx = Izz. The body frame of the quadrotor by angular velocities related to the inertial frame by\u03b3\u0307\u03c5\u0307 \u03c8\u0307  = 1 s\u03b3t\u03c5 c\u03b3t\u03c5 0 c\u03b3 \u2212s\u03b3 0 s\u03b3/c\u03c5 c\u03b3/c\u03c5 \u03c9x\u03c9y \u03c9z  , (3) where \u03b3, \u03c5, \u03c8 are Euler angles in the inertial frame, \u03c9x, \u03c9y , \u03c9z are angular velocities in the body frame, tx = tanx, cx = cosx, sx = sinx"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002113_0954407020984668-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002113_0954407020984668-Figure6-1.png",
        "caption": "Figure 6. Representation of the MBS model of the reference theoretical one.",
        "texts": [
            " In fact, this contribution is a force, applied in parallel to the system, and not a further degree of freedom and, therefore, not subjected to a possible reduction of the degree of freedoms of the system itself. The theoretical model, shown in Figure 5, has five degrees of freedom and requires the definition of 31 parameters such as inertias lengths, stiffness, damping, etc. (Table 1). Its numerical implementation was conducted in a multibody simulation environment, considered as reference for the automotive industry, MSC.ADAMS/View. The multibody model is shown in Figure 6. In Table 1 the parameters, necessary to represents the multibody model, for a total of 31 variables, are summarized. In particular, in this table the values are related to a real steering system that will be used as a test case in the paper. As concerns steering wheel, the proposed model is able to consider it into Jc parameter that represents the inertia moment that condenses column and steering wheel inertias but, in this paper the steering model does not consider steering wheel as is usually done in experimental approach to steering characterization at bench in which it is not mounted into steering line"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003416_tmag.2020.3020126-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003416_tmag.2020.3020126-Figure5-1.png",
        "caption": "Fig. 5. Prototype and testbed. (a) Laminated rotor. (b) Laminated stator. (c) PMs. (d) Assembled rotor. (d) Assembled stator. (f) Controller and inverters. (g) Testbed.",
        "texts": [
            "html for more information. However, these odd harmonics of the armature field can also generate few torques which are not considered in the MMF model. Despite the decrease, even harmonics especially the dominant harmonics in the airgap contribute to most of the total torque while the total torque can be regulated by enhancing or weakening the even harmonics. V. EXPERIMENTAL VERIFICATION To validate the previous analysis, the prototype of the proposed machine is designed and constructed as shown in Fig. 5. To ease the manufacturing process, the prototype is scaled down with relevant design parameters listed in Table VI. Due to the limited experimental conditions, each harmonic component in the air gap generated by the PMs cannot be extracted for the torque production on the prototype. Thus, the torque contribution to the average torque by each harmonic cannot be measured on the prototype. However, the FEA predicted no-load back EMFs under different DC-field currents can be verified by measurements, while the average torque generated by all the air-gap harmonics are measured and compared with the FEA-predicted one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001470_2050-7038.12588-Figure20-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001470_2050-7038.12588-Figure20-1.png",
        "caption": "FIGURE 20 Cross-sectional view of the generators: A, 48 slot; B, 60 slot; C, 72 slot; and D, 84 slot (initial)",
        "texts": [
            " Since the number and structure of the slot change the magnetic circuit of the generator, the performance of the generator is also affected.53 In this section, analysis of PMSG with various pole/slot number combination has been performed. When the number of poles is taken as fixed 14 to allow a balanced three-phase winding, the number of slot can be calculated according to Equation (11). S GCD 2P,S\u00f0 \u00de =3k \u00f011\u00de where, k = 1, 2, 3, \u2026, 2P is the number of poles, and GCD is the greatest common divisor. Slot numbers have been determined as 48, 60, 72 and transient analyzes were performed. Figure 20 provides the cross-sectional views of the PMSGs subject to the study. In the designs of the generators, physical parameters other than the slot structures were kept constant. For each different combination, the slot structure is rearranged to ensure that the stator teeth are not saturated. Performance of generators obtained in the analyzes are presented in Table 2. When the designs are examined, it is seen that the efficiency, the flux density of the stator teeth and the slot fill factor are approximately the same for each generator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure4-1.png",
        "caption": "Fig. 4. HTS-excitation coil structures: (a) Ring-shaped and (b) Racetrack-shaped.",
        "texts": [
            " As the rotor moves further 180 degrees to 270 electrical degrees, the flux linkage of coil A1 will be at its negative peak as shown in Fig. 3(b). It indicates that in the proposed RPS-HTSFS machine, the total flux linkage of the phase A by connecting the armature coils in series can be bipolar. Thus, as the rotational movement passes one by one pole pitch, the electromotive force (EMF) of the armature winging changes periodically as well. For the HTS-excitation coil, there are two feasible winding methods, namely, ring-shaped and racetrack-shaped structures as depicted in Fig. 4. In the conventional HTS-FS machine as shown in Fig. 1, several racetrack-shaped HTS-excitation coils are employed, which brings longer end winding. On the other hand, in the winding process of HTS-excitation coil, the coil needs to be applied winding tension and then the bending stress is produced. When the winding tension is over large, the inner layer of the HTS-excitation coil will be excessive bending, causing the degradation of mechanical properties. Therefore, it is of great necessity for the HTS-excitation coil to consider the proper bend radius and try to reduce its bending stress"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure4-1.png",
        "caption": "Fig. 4. The Mesh quality at high temperature.",
        "texts": [],
        "surrounding_texts": [
            "The temperature distribution on the base alloy at 293 K is shown in Fig. 2. While comparison of mesh quality plots is indicated in Figs. 3 and 4. This was done at room temperature in order to pre-heat the base metal and check the stresses via modelling. Pre-heat treatment of base alloy is good because it allows it to absorb the laser beam uniformly and avoid residual stresses that can tamper with the microstructures with larger grain sizes and influence the surface quality of the additive manufactured component. In addendum, component load resistance can be tampered with when the residual stresses are larger. Standard quality requirements must be met, and permissible irregularities must be identified and categorized. Different kind of anomalies and their influence on the additive manufactured surface performance and quality must be well understood. The wrong combinations of power, thickness of layer, direction of build, scan velocity and rate of flow can impart on the anomalies. The chief challenges debarring the full implementation of additive manufacturing is quality assurance. The real quality assurance involves predictive modelling and real time techniques of detecting anomalies immediately in additive manufacturing. Thermal cycle in additive manufacturing of metals impart residual stresses in additive manufactured parts and this is a real concern that can lead to parts distortion. The flow of molten liquid in the melt pool that came about as a result of the increase in the speed of the coating materials was linked to higher laser power. Standard distribution, enhancement and spheroidization are linked to higher laser power as shown in Figs. 5\u20137. Movement of laser from the starting point to the end changes the temperature distribution and enhanced microstructures in the melt pool are depended on the laser input, scan velocity and the faster cooling rate. The base alloy acts as heat absorber (heat sink) and the temperature gradient between the base alloy (substrate) and the incoming laser has significant effects on the microstructures. Else, larger residual stresses may be infused in the microstructures and impart on the surface quality of the com- on the base alloy at 293 K. posite coatings. The spot diameter and the distance between the base alloy and the laser nozzle influences the peak temperature distribution in the molten pool. The peak temperature at the start was 529 K while at the end of the base alloy it increased to 534 K. The crystal structures in the melt pool are changed by the rate of solidification as shown in Figs. 8 and 9. The increase in the scanning speed of the LMD operation produced less dense composites, there was not enough time for the coating materials to interact with the Laser and fuse with the base alloy. The increase in the scanning speed of the LMD operation decreased the deposit width, while the decrease in scanning speed increases the deposit width; resulting from the available time for the coating materials to interact with the laser and fuse with the base alloy [34,35]. From the temperature distribution results, it is evident that the increase in the scanning speed of the LMD operation decreases the rate of dilution of the coating materials within the structural matrix of the base alloy [35]. The propagation of these grains were initiated by the solidification process of the microstructure after the LMD operation. The laser interacts with the powders (coating materials) fed onto the surface via the feed nozzle, the laser melts the powder coatings before they arrive on the surface of the base alloy (Ti-6Al-4 V) creating a melt pool of the deposited materials directly on the surface of the base alloy. The changes in the temperature shows the interaction in the melt pool between the laser and the coating materials as the laser moves from its original position towards the end of the base alloy (substrate). Only the region of interaction between the melt pool of the coating materials and the base alloy undergo a microstructural evolution. Hence, the rate of dilution of the composites was required to understand the level of diffusion of the coating materials in the structural matrix of the base alloy. The alloying metals Al and Cu formed dendrites after the solidification process of the composites was initiated as shown in Figs. 8 and 9. Thermal gradient is the major factor that determines if dendrites would be formed in a microstructure. The alloying metal Cu from microstructural studies is known to possesses strong Beta (b) stabilizing particles that are propagated within the microstructural matrix of composites [36\u201338]. Hence the fusion of the Cu alloying element with the Titanium structural matrix initiates the propagation of the Beta-Titanium (b-Ti) phase structure. The Binary alloy Al-Cu within the structural lattice typically propagate dendritic grain structures at increased laser power and scanning speed [39\u201341]. The combination of the alpha phase particles and the beta phase stabilizers form a network of intermetallic inter-dendritic eutectic phase structure within the fusion regions of the composites [42\u201344]."
        ]
    },
    {
        "image_filename": "designv11_63_0002009_tro.2020.3033721-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002009_tro.2020.3033721-Figure3-1.png",
        "caption": "Fig. 3. Evolution of the contact frame C for a draping process along a curved path \u03c0 on a doubly-curved surface \u03c3.",
        "texts": [
            " However, in the tape application process of curved tapes, this choice for the contact frame causes wrinkles and distortions in the applied material, as the path tangent vector \u03c0\u2032(p) exhibits turns around the surface normal vector. To this end, a new frame is introduced in the next section, which solves these limitations for tape application. In this section, the aforementioned relations are extended to curved paths. The geometric relationships of a draping motion along a curved path \u03c0 on a doubly-curved surface \u03c3 are illustrated in Fig. 3 for multiple path positions p0 < p1 < p2. Instead of rotating the draping roll around the normal vector (blue), which may cause wrinkles and distortions of the tape, the roll is tilted around a particular surface tangent vector (red). The corresponding surface tangent vector field is derived as a parallel vector field of the parametrized surface \u03c3(s), starting from an initial contact frame C(p0). From this parallel vector field and the surface normal vector field, a novel parallel contact frame is constructed",
            " The tape application path is traversed with a suitable trajectory for \u03bed1 . The desired trajectory \u03bed3 specifies the draping roll motion along the surface normal vector and allows to approach the 3-D object, establish and release the contact with the 3-D object. Additionally, the tape application normal force is adjusted using the impedance parameters md 3 , kd3 , and dd3 of the coordinate \u03be3. Finally, the draping roll axis is\u2014in the absence of control errors\u2014aligned to the e2-axis of the contact frame C(p\u2217), see Fig. 3. Thus, choosing \u03bed2 = \u2212\u03c0l(p \u2217) for the lateral motion of the draping roll (see Fig. 4) correctly takes into account the lateral movement of the contact point and prevents a lateral sliding motion of the draping roll. Instead, the roll is tilted parallel to the e2-axis of the contact frame and the contact point traverses in the lateral direction appropriately. The 1-D nullspace of the redundant robot, i.e. the elbow position, is stabilized using a complementary projection matrix and a simple proportional-derivative controller [24], reading as vn = ( I\u2212 J\u0302\u2020(q)J\u0302(q) ) (\u2212Dnq\u0307\u2212Kn(q\u2212 q0)) , (36) with the positive definite controller gain matricesKn andDn and a virtual equilibrium joint positionq0",
            " The surface normal contact force f\u0302e,3 is approximately 3N, while the other forces and torques remain close to zero. 2) Impedance-based tape application: In the second phase, the tape application is performed using the surface-based path following controller (33), (35), and (36). In path coordinates y\u0302, the contact point moves from \u03be1 = 0 mm to \u03be1 = 322 mm with a velocity of \u03be\u03071 = 30 mm/s using a C2-trajectory. The lateral motion along the draping roll axis is given by \u03bed2 = \u2212\u03c0l(p), which correctly takes into account the lateral movement of the contact point, as illustrated in Fig. 8, cf. Fig. 3. While the control errors e\u03021 and e\u03022 of the geodesic direction \u03be1 and the lateral direction \u03be2 stay below 1.2 mm, the controller adjusts the position \u03be3 along the surface normal according to the impedance model. Thus, the control strategy is able to comply with uncertainties related to the 3-D object and with the position errors of the robot manipulator. Throughout the process, the surface normal force f\u0302e,3 remains the dominant component. While the surface tangent forces f\u0302e,1 and f\u0302e,2 are mostly below 1 N, the estimated torques f\u0302e,4, f\u0302e,5, and f\u0302e,6 stay well below 150 mN m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure44.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure44.3-1.png",
        "caption": "Fig. 44.3 Concept development sketches",
        "texts": [
            " The initial design encompasses a basic structure welded with a basket, which can hold beverage glasses. Gradually with a demand for stability, the gimbal mechanism is introduced, and the attributed three-point claw firm locking system brings trust and confidence. The process has gone through ample material options for manufacturing and ends up with structural strengthening design with the plastic itself. It is important to accommodate all types of glass and cups, which it does with an adjustable base (Fig. 44.3; Table 44.2). This journeywas started in order to cover all the facets of the design considerations, and concept 4was foundmost compatiblewith all aspects of the design consideration; hence, the journey comes to an end with the selection of the concept 4. Design enfolds sturdy cadmodeling for assurance and confidence. Finishedmanufacturing includes compressed injection molding with a polycarbonate plastic material ensuring reliability (Fig. 44.4). Parts 1. Basket\u2014A self-adjusting designer bottom to accommodate a range of drinking glasses and cups"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001943_012024-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001943_012024-Figure2-1.png",
        "caption": "Figure 2. Virtual test stand: (a) \u2013 different schemes of a machine-tractor unit; (b) \u2013 various types of obstacles.",
        "texts": [
            " Also, additionally, a spherical element was installed to compensate for the mass-inertial characteristics of the non-simulated stationary elements of the tractor. He has the ability to change the spatial position and mass. By adjusting these parameters, the mass-inertial characteristics of the tractor model were brought to the parameters of a real tractor. The tractor model was equipped with front and rear mounted forestry modular tillage implements [4] (figure 1). The following implements were used in the modeling: front-mounted drum chopper (figure 1a), rear-mounted single-row harrow (figure 1b); rear-mounted two-row harrow (figure 1c). Figure 2a shows three typical unit layout options: unit with a single-row rear-mounted implement (scheme Forestry 2020 IOP Conf. Series: Earth and Environmental Science 595 (2020) 012024 IOP Publishing doi:10.1088/1755-1315/595/1/012024 0 + 1), two-row rear-mounted implement (scheme 0 + 2), and with front-mounted and two-row rear-mounted implements (scheme 1 + 2). Depending on the applied reforestation technology and operating conditions, the implement s can be equipped with sets of various working tools",
            " The overall dimensions of the working tools and their massinertial characteristics are maximally unified. As a result, the mass-inertial characteristics of modular implements for various purposes are also very similar. In the study, a front-mounted drum chopper, a rear-mounted one and two-row harrow was used to complete the unit. Four test sites were created with semi-cylindrical immovable obstacles located on them, oriented perpendicularly to the direction of movement of the tractor, which simulates movement in forest areas (figure 2b), these are: single linear (1); single sequential (2); group linear (3); and group sequential (4). The height of single obstacles 1 and 2 was 200 mm, group obstacles 3 and 4 was 150 mm. To speed up the modeling process, all four types of obstacles are combined into a single test track 40 meters long and 6 meters wide. The interaction between the contact pairs \u201csupport surface \u2013 tractor wheels\u201d had the parameters of the standard \u201csteel \u2013 rubber\u201d interaction taken from the SolidWorks Motion library, but the coefficients of dynamic and static friction were modified"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000360_s00158-019-02415-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000360_s00158-019-02415-3-Figure1-1.png",
        "caption": "Fig. 1 Bucket-wheel stacker reclaimer",
        "texts": [
            "eywords Weight reduction optimization . Forearm of bucket-wheel stacker reclaimer . Surrogate model . Morris method . Sequential multi-point infill criterion As a bucket-wheel stacker reclaimer, shown in Fig. 1, combines stacking, reclaiming, and transporting with the belt conveyor system, it has advantages of high efficiency, simple operation, safety, and reliability. Nowadays, it is widely used in the places such as large open-pit mine storage yards, power plants, gas plants, large ports as well as the fields like power, transportation, mining, metallurgy, building material, chemistry, etc (Yang et al. 2015; Wang 2012; Wu et al. 2014). The productivity of a bucket-wheel stacker reclaimer is 1.5~2.5 times that of a single-bucket loader (Shao et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure4.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure4.2-1.png",
        "caption": "Fig. 4.2 A layout of the differential hydraulic cylinder",
        "texts": [],
        "surrounding_texts": [
            "observer is provided to both loops. It is noted that the load force in the outer piston position control loop is not reconstructed through the estimator."
        ]
    },
    {
        "image_filename": "designv11_63_0001589_022060-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001589_022060-Figure6-1.png",
        "caption": "Figure 6: Schematic illustration of protruding poles (left) and static eccentricity (right).",
        "texts": [
            " The force densities can be derived from the well-known Maxwell stress tensor and are calculated where \ud835\udc35rad and \ud835\udc35tan are the magnetic flux density components in radial and tangential direction respectively depending on time t and circumferential coordinate x: \ud835\udf0erad(\ud835\udc65, \ud835\udc61) = 1 2\u03bc0 (\ud835\udc35rad 2 (\ud835\udc65, \ud835\udc61) \u2212 \ud835\udc35tan 2 (\ud835\udc65, \ud835\udc61)) \u2248 1 2\u03bc0 \ud835\udc35rad 2 (\ud835\udc65, \ud835\udc61), (1) \ud835\udf0etan(\ud835\udc65, \ud835\udc61) = 1 \u03bc0 \ud835\udc35rad(\ud835\udc65, \ud835\udc61) \ud835\udc35tan(\ud835\udc65, \ud835\udc61). (2) In (2), it is assumed that \ud835\udc35rad \u226b \ud835\udc35tan. The force excitation of the structural elements can be calculated by surface integrals over the structural areas adjacent to the air gap of the generator. In the case of the stator, the teeth of the electrical steel core conducting the magnetic flux exert these forces. Air gap imperfections resulting from wind loads, inhomogeneous thermal conditions or manufacturing tolerances of the machine will change the local magnetic field distribution. In Figure 6, schematic illustrations of a protruding pole and a static eccentricity are shown. Both effects lead to a deviation \ud835\udf16(\ud835\udf11) from the ideal air gap width \ud835\udeff0 and therefore to the air gap width distribution \ud835\udeff(\ud835\udf11), depending on the circumferential position determined by the angle \ud835\udf11: \ud835\udf16(\ud835\udf11) = \ud835\udeff0 \u2212 \ud835\udeff(\ud835\udf11). (3) Predictive models must take these effects into account, since they have a significant impact on the electromagnetically excited forces [8]. The air gap width distribution represents accordingly one dimension of the shown look-up table-based method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure7.4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure7.4-1.png",
        "caption": "Fig. 7.4 (a) Wire components used in the arc spray process and (b) principle of the wire arc spray technique",
        "texts": [
            " During this process, the substrate receives heat from the solidifying splats, but it is also heated by the tail end of the flame, as the hot gas jet escapes from the impact area by flowing out sideways over the substrate. These two sources of heat can significantly raise the temperature of the substrate and may affect the properties of the temperature-sensitive light alloy. Al-SiC composites [7] and Ni-SiB-Ag alloys [14] are materials that have been deposited on light alloys by this technique. In the twin arc spray process, the material precursor is fed into the system in the form of two cored wires, which melt within an electric arc. Figure\u00a07.4a schematically shows the wire preparation to be used in this process. In the case of composite wire, it is formed by a metal alloy sheath and a powder that is placed inside such a\u00a0sheath. The powder can be contain different particles sizes or complex chemical compositions, such as FeCrBSiNbW [9], FeBSiNbCr, and FeBSiCrNbMnY, which have high glass-forming ability, obtaining a composed coating microstructure formed by nanocrystalline phases in an amorphous matrix [15]. Compared with 7 Thermal Spray Coatings 109 other thermal spray processes, this process has some advantages due to its low cost and easy application. Further, the coatings can be synthesized successfully. Figure\u00a07.4b shows the fundamentals to obtain a coating using the arc wire spray technique. The electrical arc spray process uses a direct current electric arc, struck between two consumable electrode wires, to melt the wires. A high-pressure gas jet is used to atomize the melts on the tips of the wires to create droplets with high velocity. The high spray rate of the\u00a0arc process for different materials makes it the most cost-effective process. Arc spray is available to use with\u00a0conductive filler\u00a0materials in the form of wire, powder or powder-filled wire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure2-1.png",
        "caption": "FIGURE 2. Boundary Condition",
        "texts": [
            " Three type of material were used in the simulation, namely: JIS 3445 STKM 13A, Aluminum 6005A T6 and Aluminum 6061 T6. The properties of material were shown in Table 2. 030153-2 Boundary Condition (BC) must be set on the model before simulation. BC must represent the real condition of object. In this paper BC was lied on the bottom part of frame since this part was connected to the chassis of electric bus. The type of BC is fixed which the rotation and displacement are not allowed in all directions. The location of BC was marked by the red color on Figure 2. While the load with the magnitude and angle based on the ECE 66 standard was shown in Figure 3. 030153-3 The output of simulation is static deflection distribution along the frame that shown in Figure 4. From the results of static simulation, it can be seen the value of static deflection on the structure of the electric bus frame. The maximum deflection value is 756,69 mm while for the minimum deflection it is 84.06 mm and the average static deflection value is 420.03 mm. After the static deflection value was obtained from the static loading simulation results, then the impact factor values are calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.6-1.png",
        "caption": "Fig. 34.6 Detailed structural design (left) and detailed surface design (right)",
        "texts": [
            " Further, a fixture provided below the horizontal stabilizer facilitates easy attachment of the boom and empennage. Horizontal stabilizer has a negative angle of incidence of 6.5\u00b0 to balance the overall moment of 4554 N-mm of the aircraft during the cruise. All structures were given sufficient tolerances ranging from 0.5 mm to 2 mm as per the need for the structure, to make sure that we have no hindrances while fabricating and as a result we had zero clashes between the parts in Catia assembly. A covering film is added on the structure, which acts as a skin for the air-craft. Figure 34.6 shows the structural design and the surface design of the UAV. The wing and tail is assembled with the fuselage by using telescoping method, and their freedom is arrested using nuts and bolts. Fully assembled aircraft has the overall dimensions as listed in Table 34.10. Considering the wing as a cantilever beam a structural analysis was performed by adding a load of 120 N (calculated using ultimate load factor) on the tip of the wing and fixing the root of the wing showed that the maximum stress was on the carbon fibre spar"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure6-1.png",
        "caption": "Fig. 6. CFD simulation model",
        "texts": [
            " The copper loss assigned to the slot node is the total copper loss of the machine divided by the number of slots, and the core loss assigned to the yoke node is \u03b8 2\u03c0Pcore where \u03b8 is the arc angle of the yoke considered in the LPTN and Pcore is the total core loss in the machine. A CFD analysis is done with controlled flow rates ranging from 2-16 liter/minute to accurately determine the temperature distribution and temperature rise in different sections in the liquid-cooled stator. To reduce computational complexity, only a section of the stator consisting of one slot and two winding supports and the portion of the stator yoke connected to these components are considered as presented in Fig. 6. The copper and core losses are introduced in the simulation as the internal heat source of the winding and yoke, respectively. Moreover, to reduce the pressure drop or reduce the coolant path, four inlets and four outlets are used in this analysis. The considered coolant in this research is water-ethylene glycol (50-50%) mixture as this is the widely used coolant for traction 1085 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 19,2021 at 14:42:00 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001894_icem49940.2020.9270713-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001894_icem49940.2020.9270713-Figure1-1.png",
        "caption": "Fig. 1. Motor geometry (1/8 model)",
        "texts": [
            " In this paper, FEA based efficiency maps are made with different loss calculation models for an IPM machine and compared with a measurement result to obtain appropriate fidelity for the FEA calculations. After observing loss maps, the contribution of loss components to the efficiency map and presents the causes of the loss components are discussed. To validate the accuracy of the efficiency map simulation, the measurement results of a 3-phase, 8-poles, IPM machine are used. A partial 3D (1/8) of the motor is shown in Fig.1. The specification of the motor is given in Table 1. This motor is used as the traction motor of an electric vehicle. The motor is driven by the PWM inverter with a carrier frequency of 12 kHz and the input voltage specified using current vector control. The diagram of the measurement system is shown in Fig.2. The speed of the motor is controlled T Hiroyuki Sano with JSOL Corporation, JMAG Division, Tokyo, Japan (e-mail: sano.hiroyuki@jsol.co.jp) Nicolas Schneider with JSOL Corporation, JMAG Division, Tokyo, Japan (e-mail: nicolas"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001671_icuas48674.2020.9213853-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001671_icuas48674.2020.9213853-Figure2-1.png",
        "caption": "Fig. 2: Forces acting on the quadrotor and the slung load system.",
        "texts": [
            " This crucial assumption decouples the rotational dynamics of the quadrotor from the rotational motion of the slung load. 1In the following, the notation sin(x) = sx, cos(x) = cx and tan(x) = tx for any angle x is considered. 323 Authorized licensed use limited to: University of New South Wales. Downloaded on October 18,2020 at 08:29:36 UTC from IEEE Xplore. Restrictions apply. The rotational dynamics of the slung load \u03b8p is developed from the fact that the vectors lp, T12, and T21 are colinear, see Figure 2, as a consequence, lp\u00d7 T12 = 0, (10) lp\u00d7 T21 = 0. (11) Replacing (4) into (10) and taking into account equation (7), one obtains mQlp\u00d7 p\u0308 = TTp\u00d7Reb3. (12) Being p a unitary vector, it satisfies the following identities p>p = 1, p\u0307>p = 0, p>p\u0308 = \u2212p\u0307>p\u0307. Now, from the triple vectorial product, one has that p\u00d7 (p\u00d7 p\u0308) = p(p>p\u0308)\u2212 p\u0308(p>p). As a result, p\u0308 takes the form p\u0308 = \u2212p(p\u0307>p\u0307)\u2212 p\u00d7 (p\u00d7 p\u0308). (13) Replacing p\u0308 from (13) into (12) results in (mQ +mp)(r\u0308p \u2212 gee3) = \u2212TTReb3 \u2212mQlp(p\u0307>p\u0307) \u2212p\u00d7 (mQl(p\u00d7 p\u0308))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001299_s00202-020-01062-y-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001299_s00202-020-01062-y-Figure10-1.png",
        "caption": "Fig. 10 Canned motor designed for tests",
        "texts": [
            " B is another constant characteristic of the insulation population, test specimen, test method, and failure mode. The coefficients A and B are calculated by fitting the above equation to experimental data. This fitting is done by the method of least squares. The lifetime curve of the tested insulation material is illustrated in Fig.\u00a09. To validate the developed electrical and thermal models, a canned motor was designed and instrumented with temperature sensors in the stator coil heads, in the slots, on the motor housing and at the rear bearing as shown in Fig.\u00a010. An instrumented test bench has also been designed. It is composed of the motor to be tested, a tank containing the fluid that cools the motor and lubricates the bearings, temperature sensors, manometers, a flowmeter, a valve, and recirculation pump as shown in Fig.\u00a011. (34)(X) = 1 T , (35)B = E 2.303Rgas . In order to validate the model, several tests were conducted on an instrumented canned motor that allows identifying the mechanical losses and all the electrical parameters for the motor performance analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure4-1.png",
        "caption": "Fig. 4. Open-circuit field distribution of the CPM-I. (a) Flux lines. (b) Flux density.",
        "texts": [
            " To ensure a good flux-modulation effect, the PM pole pairs are always large [24]. Since the optimal \u03b1 is 0.66, the iron pole is relatively thin. This means that the flux density in salient iron pole is high. Together with elevated armature flux due to reduced effective airgap length, the CP rotor is prone to magnetic saturation. This could limit their overload torque capability. All the radially magnetized PMs are homopolar in the CP PM machine. As a result, the flux lines pass through the iron poles radially, as shown in Fig. 4 (a). Due to the fringe leakage flux, the flux density in Region-II is larger than that in Region-I, resulting in severer saturation in Region-II, as shown in Fig. 4 (b). However, if larger iron pole width is adopted to reduce the saturation, the electromagnetic performance will be decreased as well. Hence, it is necessary to find a way that can reduce the flux density of the iron pole while maintaining the Authorized licensed use limited to: Rutgers University. Downloaded on May 20,2021 at 12:02:13 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002881_s42417-021-00329-3-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002881_s42417-021-00329-3-Figure19-1.png",
        "caption": "Fig. 19 Isotherms diagram of box wall",
        "texts": [
            " The temperature measurement point in the experiment was the outer end of the bearing housing (10\u00a0mm away from the heat source position), and during simulation calculation, if the temperature was extracted directly at the heat source, the flash temperature phenomenon at the heat source would affect the temperature extraction results. Therefore, under the premise of considering the effect of heat conduction and the assumption of material isotropy, the location with the same distance between the measuring point and the heat source was used as the temperature extraction point of the simulation calculation results. Figure\u00a019 shows the diagram of the steady-state isotherms of the outer surface of the gearbox. The surface temperature distribution shows a pattern of uneven diffusion from the position of bearing 1 with the largest heat generation to the surroundings. The temperature around bearing 3 and bearing 5 shows a tendency to gradually decrease as the distance from bearing 1 increases. The position of bearing 1 with the highest speed then becomes the area with the highest temperature, reaching 13.5\u00a0\u00b0C, while the increase 1 3 in temperature of bearing 3 and bearing 5 was 51"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure6.10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure6.10-1.png",
        "caption": "Fig. 6.10 Schematic representation of the remanufacturing process carried out on a metal piece. (a\u2013e) A metal coat is deposited along the bar to restore functional working dimensions",
        "texts": [
            " It is due to this last characteristic that cold spray has been widely used by the industrial sector. This technology is being applied in remanufacturing operations of metal components where operating temperatures below the melting temperature are required. Thus, it is ideal for depositing materials in which it is desired to retain a large part of their initial properties, as well as producing oxide-free deposits without an adverse influence on the substrate surface on which the deposit has been made (Fig.\u00a06.10). In the manipulation of light alloys, such as those of aluminum and magnesium, low-pressure cold spray systems are widely used to carry out this kind of deposits, with the aim of applying a protective layer against corrosion on chemically compatible substrates. Cold spray is a technology characterized by the absence of high temperatures in its process. In such a process, small particles (~150\u00a0\u03bcm) are accelerated at a supersonic speed through the effect of a pressurized gas, which can be nitrogen or helium, and whose operating temperature reaches 800\u00a0\u00b0C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure10.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure10.2-1.png",
        "caption": "Fig. 10.2 A single-phase model of LCL output filter",
        "texts": [
            " The shunt active filter is generally composed of a control part and a power part, including a PWM voltage source inverter (VSI), a DC capacitive storage system, and an output filter. The control part usually consists of the current perturbation identification block and the control loops for injecting currents into the grid and for regulating the direct voltage storage capacitor. filter out the components of the switching frequency, generated by the PWM-VSI. A single-phase equivalent circuit of the LCL output filter is shown in Fig. 10.2. The LCL filter is modeled by the following Laplace domain equations [4]: Iinj = B1(s) A(s) Vf (s) + B2(s) A(s) Vs(s) (10.1) with \u23a7 \u23a8 \u23a9 A(s) = a1s3 + a2s2 + a3s+a4 B1(s) = b11s + b12 B2(s) = \u2212(b21s2 + b22s + b23) and a1 = (Ls + Lf 2)Lf 1Cf a2 = (Ls + Lf 2)Rf 1Cf + (Rs + Rf 2)Lf 1Cf + (Lf 1 + Ls + Lf 2)Rf Cf a3 = Ls + Lf 2 + Lf 1 + (Rs + Rf 2)Rf 1Cf + (Rf 1 + Rs + Rf 2)Rf Cf a4 = Rf 1 + Rs + Rf 2 b11 = Rf Cf b12 = 1 b21 = Lf 1Cf b22 = (Rf + Rf 1)Cf b23 = 1, where Lf 1 is the inverter side inductor, Lf 2 is the grid-side inductor, Cf is a capacitor with a series Rf damping resistor, Rf 1 and Rf 2 are the inductor resistances, Ls and Rs are grid inductor and resistor, respectively, B1(s) A(s) is the transfer function model of the LCL filter (including the network impedance), B2(s) A)(s) represents a disturbance model, Vf (s) is the inverter output voltage, and Vs(s) is the point of common coupling (PCC) network voltage, which is considered as a bounded disturbance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001894_icem49940.2020.9270713-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001894_icem49940.2020.9270713-Figure15-1.png",
        "caption": "Fig. 15. Comparison of iron loss density distribution",
        "texts": [
            " To consider the effect of the stress, a stress analysis is conducted to obtain the stress distribution in the machine as shown in Fig. 13. A high-stress distribution is observed in the root of the stator tooth where the change in the material\u2019s magnetic property will be the largest. In the magnetic analysis, a stress-dependent iron loss property is used as shown in Fig. 14. The increased ratio of the iron loss is larger for low magnetic fields and this is the reason why the effect of the stress is the highest at the low load region. The resulting iron loss density distribution is shown in Fig. 15. The accuracy of the FEA-based efficiency map was evaluated, comparing it to measurement results for an IPM Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 24,2020 at 06:42:49 UTC from IEEE Xplore. Restrictions apply. machine. It was found that high fidelity loss modelling is crucial to achieving reasonable accuracy with an error of less than 1%. Using high-fidelity loss modelling, it is possible to observe how the loss occurs at various operating points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002411_iros45743.2020.9340935-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002411_iros45743.2020.9340935-Figure8-1.png",
        "caption": "Fig. 8: (a) The rotation angle of the inchworm varies within two fixed angles \u03b11 and \u03b12. (b) Velocity of the inchworm robot measured in different test conditions. (c) Definition of the turning radius . (d) CAD model of the turning inchworm robot. The inchworm robot turns to the right as the front servo is actuated.",
        "texts": [
            " Three main tests were performed: individual inchworms assembling into a quadruped robot which stands on four legs and walks; inchworms passing through confined spaces; and the quadruped robot climbing obstacles. We first demonstrate the locomotion of a single inchworm robot. In BioARS, the time required for gathering and assembling is determined by how quickly and nimbly the modules can move. To transfer the buckling motion of the inchworm robot into crawl locomotion, the inchworm robot was set on a platform covered with nylon cloth. The inchworm robot rotates between two angles, the minimum rotation angle \u03b11 and the maximum rotation angle \u03b12 (Fig. 8(a)). The maximum rotation angle \u03b12 is set as 120\u00b0, 150 \u00b0, and 180\u00b0. The frequency of buckling motion f is set as 1 Hz and 2 Hz. The minimum rotation \u03b11 ranges from 90\u00b0 to 180\u00b0. The velocity of the inchworm robot is recorded in Fig. 8 (b). The turning test of the inchworm robot mainly focused on the turning radius (Fig. 8(c)). The buckling motion of the inchworm robot is driven by the rotation of the pair of servo motors and through pulling the cables. During the turning, we only actuate one servo to pull the cable. Thus, the inchworm robot will buckle asymmetrically (Fig. 8(d)). The experiment results show that the minimum turning radius is about 80cm while the actuated servo is set as \u03b11 = 90\u00b0. Previous work has proved that the \u201chead-to-head\u201d connection is capable of resisting tensile load and shearing load [18]. However, this connecting method cannot meet the requirement of the high bending moment resistance in our model. As the quadruped robot stands and walks, the joints of the quadruped robot are under large bending force, and the failure bending moment of the magnetic connection is relatively large"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000052_s1068366619050210-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000052_s1068366619050210-Figure2-1.png",
        "caption": "Fig. 2. Power and geometric relationships: a is when tightened in the free state, and b is with an external load.",
        "texts": [
            " Here: and are the bearing dimensions, and are the diameters at the bottom of the groove of the inner and outer rings, and are the diameters of the sides of the base ends of the rings, is the diameter of the lock on the bottom end of the outer ring, and is the ball diameter. d D id oD bd bD lD wD 425 When the outer rings are tightened, all the balls are loaded evenly and the contact angles with the inner and outer rings will be the same. In this case, the force and geometrical relations for an individual ball, without taking into account deformations of the rings during planting and thermal expansion, will be determined for each bearing, as shown in Fig. 2a. Here: is the axial displacement of the centers of the gutters in the section, and is the distance between the centers of the grooves of the rings. In addition, we have: where is the total deformation in contacts and is the displacement of the considered ring during compression of the rings. The angle of contact of the ball with the grooves of the rings and the tightening force are determined by the formulas: where is the number of rolling elements in the bearing. The coefficients of compliance of contacts, taking into account the previously given factors, are determined by the formulas: (1) In this case, the radii of curvature in the contacts have the form (2) tight1 1 0 iA\u0421 S S= = + \u0394 1 1 0B C H= tight tight 2 2 1 1 0 ,mA B S H R= + = + \u03b4 ( ) ( )( )\u03b4 = \u03bb + \u03bb \u03b1 2 3 tight i o tight tightsinF z i\u0394 ( ) ( ) ( )( ) \u03b1 = = \u03b1 + \u2212 \u03bb + \u03bb tight tight 0 tight 1",
            " In the presence of initial axial play and loading of the bearing control axial force there will be an axial displacement of the rings. In this case, the load will be perceived by only one row of balls. The efforts in the contacts of the loaded row will be equal: We accept that all the balls are loaded uniformly and the contact angles with the inner and outer rings will be the same. In this case, the force and geometrical relations for an individual ball will be determined for each row, as shown in Fig. 2b. In Fig. 2b: is the axial displacement of the centers of the gutters in the section, and is the distance between the centers of the grooves of the rings along the radius. tight1 tight2 1 2d .an F F\u0394 = \u0394 + \u0394 = 1 2\u0394 = \u0394 tight1 01 1S S= + \u0394 2 1\u0394 = \u0394 \u2212 \u0394 tight2 02 2S S= + \u0394 ( )tight1 tight1 01arctan S H\u03b1 = ( )tight2 tight2 02arctan S H\u03b1 = \u03b4ik \u03b4ok riR roR \u03bbi \u03bbo tight tigh 2 2 1 1t 01 1mS H R\u03b4 = + \u2212 tight tigh 2 2 2 2t 02 2mS H R\u03b4 = + \u2212 ( )( )= \u03b1 \u03b4 \u03bb + \u03bb 1.5 1 1tight tight tigh i1 o1t1sinF z ( )( )= \u03b1 \u03b4 \u03bb + \u03bb 1.5 2 2tight tight tigh i2 o2t2sinF z ( )tight tig1 2ht tight1F F F\u2212 1tight tightF F= 1\u0394 0F ( )= = \u03b1o i 0 sin "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002010_j.tws.2020.107334-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002010_j.tws.2020.107334-Figure3-1.png",
        "caption": "Fig. 3. Total configuration of the screw-type actuator.",
        "texts": [
            " In order to prevent the relative motion between the actuator and the cylindrical shell, several axial sliding blocks and circumferential sliding blocks are used to limit its movement in axial and circumferential Fig. 10. The SMA driving tube. Y. Lu et al. Thin-Walled Structures 159 (2021) 107334 directions. The sliding blocks are fixed on the cylindrical shell by bolts to ensure that the screw-type actuator is closely attached to the inner wall of the shell. The total configuration of the screw-type actuator is illustrated in Fig. 3. Fig. 4 gives the schematic diagram of the screw-type actuator. One can find that the actuator is mounted inside the cylindrical shell. The driving sources of the actuator are SMA tubes with the size of \u00d89mm \u00d7 19.7 mm \u00d7 2.5 mm. The length of the SMA tube after compression in martensite state is 19.04 mm. When heated in unconstrained state, it can recover to 19.64 mm, and the recovery stroke is about 0.6 mm. Fa denotes the driving force generated by the SMA tube. When both ends of the SMA tube are restrained, a single SMA tube can generate a driving force of (0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001432_042017-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001432_042017-Figure6-1.png",
        "caption": "Figure 6. SGU structure",
        "texts": [
            " The above problem also causes high carbon emissions of computers, making artificial intelligence been questioned in terms of environmental protection. The reason is that LSTM was originally designed to handle the complex logic of natural language. But the logic of frequency domain energy feature is relatively simple. While the bearing fault diagnosis task needs higher real-time performance. Therefore, a corresponding simplified strategy is proposed, a new structure SGU with only one \"gate\" is designed as shown in figure 6, Formulas are shown in (10) (13). 1c tanh( )t xc t hc t cW x W h b       (10) 1 1 ( ) t xg t hg t cg t g g sigmoid W x W h p c b          (11) t 1c ( )t t tc c g   (12) th tanh( )tc (13) where\uff0c xcW  \uff0c hcW  \uff0c xgW \uff0c hgW NR N \uff0c cgp \uff0c cb \uff0c gb NR are the parameters that the network needs to learn. ICEMCE 2020 Journal of Physics: Conference Series 1601 (2020) 042017 IOP Publishing doi:10.1088/1742-6596/1601/4/042017 Considering that SGU can only use current and previous information, we propose bi-directional SGU (Bi-SGU) for bearing fault diagnosis to make full use of information at every time step"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000274_icems.2019.8921504-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000274_icems.2019.8921504-Figure1-1.png",
        "caption": "Fig. 1. Structure of four topologies. (a) Model 1. (b) Model 2. (c) Model 3. (d) Model 4.",
        "texts": [
            " In this paper, four PMVM topologies are proposed. With same stator and winding configuration, their permanent magnet(PM) configurations are different. Based on FEM, their electromagnetic performance, such as air-gap flux density and back electromotive force (EMF), are analyzed and compared. Additionally, in order to find the best structure parameters, the influence of parameters on performance is analyzed. II. TOPOLOGIES AND WORKING PRINCIPLE Four PMVM topologies with different permanent magnet configuration are shown and introduced in Fig.1 and they are named from model 1 to model 4. Model 1 is existing topology with traditional surface PMs. Model 2 is topology with surface halbach PM. Model 3 has interior V-type PMs and Model 4 has interior spoke-array PMs. The working principle of PMVM, magnet field modulation principle, is analyzed and validated in this section. The initial parameters of four topologies are listed in Table I. If PM pole-pair number on the rotor is Qpm, the fundamental wave of magnetomotive force (MMF) can be expressed as follow: = cos[ ( )]a pmF F Q t\u03b8 \u03c9\u2212 (1) where Fa, \u03c9, \u03b8 and t are amplitude of MMF fundamental wave, rotor angular velocity, space position and time",
            " One is the cogging torque, which is caused by the slots on the stator, and the other is the ripple of electromagnetic torque, which is caused by the harmonics of back EMF. Comparison of four topologies is shown in Table III. With initial parameters, Model 1 has advantages in most of performance indexes. IV. INFLUENCE OF STRUCTURE PARAMETERS Different structure parameters have different influence on performances of PMVM. Appropriate structure parameters can lead to higher torque and low torque ripple. In this section, several structure parameters are chosen to be analyzed. It can be seen in Fig.1 that four topologies have different PM shapes. PM of Model 1 and Model 2 are tile-shape while that of Model 3 and Mode 4 are rectangular. So in this section, different PM parameters will be analyzed on base of stationary volume of PM. In Model 2, PM are magnetized in radial or tangential direction. The total pole-arc coefficient \u03b1p is 1 and the initial pole-arc coefficient of radially magnetized PM, \u03b1pr, is 0.5. Fig.7(a) shows the average torque and torque ripple with \u03b1pr ranging from 0.1 to 0.9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002190_012039-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002190_012039-Figure2-1.png",
        "caption": "Figure 2. Design diagram of a pneumatic drive",
        "texts": [
            " The mathematical model of the pneumatic drive demonstrates a system of differential equations describing the movement of the working body and the change in pressure in the cavities of the actuator, the mathematical model of the pneumatic drive includes the following equations: 1. The equation of motion of the actuator of the pneumatic cylinder. 2. Equations of pressure change in the discharge cavity. 3. Equations of pressure change in the exhaust cavity. The design diagram of the pneumatic drive is shown in Figure 2. Dynamics of Technical Systems (DTS 2020) IOP Conf. Series: Materials Science and Engineering 1029 (2021) 012039 IOP Publishing doi:10.1088/1757-899X/1029/1/012039 When designing, one of the main conditions is to confirm the functionality of the developed drive, as well as to analyze the processes occurring in it when positioning the pneumatic cylinder. When forming the mathematical model of the drive, the following assumptions were made:  the pressure of the compressed air source is constant over time;  the thermodynamic process of gas behavior in the pneumatic system is assumed to be adiabatic;  in the description of pneumatic devices, the ideal gas model is used, since the pressure in the pneumatic system is below 10 bar;  leaks are not taken into account;  the force of viscous friction is proportional to the speed;  the coefficients of expenses are taken as averaged;  the mass of the moved parts is assumed constant;  force Fc at the output link of the pneumatic drive is constant;  relay control of pneumatic valves;  the time of forming the control signal from the displacement sensor is not taken into account [4-5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002095_012032-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002095_012032-Figure7-1.png",
        "caption": "Figure 7. Model of the laminated composite material in tensile test (a) scheme of loading (b) meshing with shell elements",
        "texts": [
            "1088/1757-899X/998/1/012032 A parameterized model with the finite element of the two laminates were constructed in Abaqus software to perform the simulation of mechanical behavior such as Stresses, strain, and displacements in both tension test and bending test. Here, the results attained by FEA in the bending test would compare with experimental results for the elastic domain only. It is shown that the composite layup tool for shell-type parts has facilitated to create stacking sequence and ply orientation as per requirement. It was designed a shell whose dimensions are 127 x 20 mm in bending and 165 x 20 mm in the tensile test. The scheme of loading & mesh model in both bending and tensile test was as shown in figure 6 and figure 7. In tensile test, Laminate-1 employed of maximum load of 4281 N with total elements about 672, Laminate-2 subjected to maximum load of 4030 N with 656 total elements and for bending test maximum force applied for elastic domain about 100N for both specimen composites. In this paper, the finite element model was defined with a 4-node doubly curved S4R element which was used for thin or thick shell element for both tests. Boundary conditions were applied according to loading for tensile and bending test separately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001273_j.istruc.2020.07.030-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001273_j.istruc.2020.07.030-Figure2-1.png",
        "caption": "Fig. 2. Basic curved beam element with section springs at ends.",
        "texts": [
            " At present, almost all existing curved beam elements only discussed the situation of rigid joints. To fill this gap, the curved beam element with one or two pinned joints are presented in this section. When an element end is assumed as a pinned joint, the moment at this end is zero but the end rotation still exists. For the displacementbased method, it is impossible to derive element formulation with pinned joints by directly using the element model in Section 2. Thus, two section-springs with zero length should be assigned in the element ends, as shown in Fig. 2. This method is often used for straight beam elements to consider semirigid joints. The external moment between the section-spring and the other element\u2019s node can be given by: = \u2212 =M S \u03b8 \u03b8 j( ), 1, 2je j je ji (13) in which Sj is the stiffness of the section-spring; \u03b8je and \u03b8ji are the rotations of the section-spring and the internal element node j, respectively. Also, according to the moment equilibrium conditions at the section-spring, the internal moment between the internal beam element and the section-spring is: = \u2212 = \u2212 \u2212 =M M S \u03b8 \u03b8 j( ), 1, 2ji je j je ji (14) In terms of the internal curved beam element between these two section-springs, the element formulation has been introduced in Sections 2 and can be given by \u23a7 \u23a8 \u23a9 \u23ab \u23ac \u23ad = \u23a7 \u23a8 \u23a9 \u23ab \u23ac \u23ad P M M k e \u03b8 \u03b8 [ ]i i e i i 1 2 1 2 (15) Assembling Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure7.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure7.2-1.png",
        "caption": "Fig. 7.2 Schematic general diagram of physical and chemical phenomena involved in the substrate/coating system obtained by thermal spray processes",
        "texts": [
            " Powder materials are directly fed into a high temperature flame through a powder feeder to complete the first stage, while with wire or rod material spray droplets are created by atomizing the melting tip as it is fed into the flame. To completely or partially melt spray materials, various heat sources including gas flames, electrical arcs, and plasma jets are employed. The heating ability of a heat source is limited by its maximum temperature, which determines the types of materials that can be applied with a specific spray process, as will be explained in detail in Sect. 7.3. Some physical and chemical phenomena are involved in the substrate/coating system, as shown in Fig.\u00a07.2. These phenomena depend on the thermal spray process used, and some of these features lead to a good or bad coating, due to the high amount of variables that each process involves. In general, thermal spray coatings 7.1 Introduction to\u00a0Thermal Spray Processes 106 can have between 100\u00a0\u03bcm and 3\u00a0mm in thickness, which is controllable depending on the property to be improved. For the thermal spray process, deposits generally have a lamellar structure if they are not posttreated. The main adhesion mechanism is mechanical, resulting in relatively low bonding strength"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000680_ab788d-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000680_ab788d-Figure5-1.png",
        "caption": "Figure 5. Schematic illustration of PEC response mechanism on the TiO2NSs/C-dots/GCE sensor.",
        "texts": [
            "2-times larger than that of C-dots/GCE and TiO2NTs/GCE, respectively, revealing that the TiO2NSs/C-dots nanocomposite presents higher photocatalytic and PEC performance than C-dots or TiO2NSs in PCP detection. The enhanced performance can be owing to the synergistic effect between 2D TiO2NSs and C-dots. The photo-generated electrons from C-dots will migrate to the conduction band (CB) of TiO2NSs across the interface, while the holes irradiated by visible light could not transfer to the more positive valence band (VB) of TiO2NSs, leading to the hindrance of electron-hole recombination, as shown in Fig. 5. When PCP molecules are captured by the TiO2NSs/C-dots surface, they could be oxidized directly on the surface, and the oxidation product of PCP could further avoid electron-hole recombination effectively, thus increasing the photocurrent signal. Optimization of detection conditions.\u2014To select suitable electrolyte, several kind of buffers including PBS buffer (0.2 M, pH 7.0), NaAc-HAc buffer (0.2 M, pH 5.0), and Tris-HCl buffer (0.1 M, pH 7.0) are investigated. Figure 6a shows that the PEC photocurrent of 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002347_j.msea.2021.140880-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002347_j.msea.2021.140880-Figure14-1.png",
        "caption": "Fig. 14. Von-Mises stress distribution for (a)\u2013(c) the overall tendency, (d)\u2013(f) the cut surface stress distribution, and (g)\u2013(i) the top surface stress distribution for three specimens.",
        "texts": [
            " Details of Ti\u20136Al\u20134V were provided for each specimen analysis. A fixed boundary condition was given on the narrow side surface for each structure, and a tensile force of 100 kN was applied on the other side for the mechanical analysis which used linear elastic model based on Hooke\u2019s law. The temperature of the bottom side was set to 1000 \u25e6C for each specimen for the thermal analysis with the top side natural heat release to the air with the convection boundary condition. The von Mises stress distribution for the tensile simulation is presented in Fig. 14. As shown in Fig. 14(a)\u2013(c), the ridge area for each specimen has a relatively low von Mises stress. It can be observed that there is a clear stress concentration in the pit area of the J. Lee et al. high-roughness specimen; however, it is relatively low in the smooth and low-roughness specimens (Fig. 14(d)\u2013(f)). This stress concentration in the pit area for the high-roughness surface can be clearly detected in Fig. 14(g)\u2013(i). Fig. 15 shows the thermal analysis results for three different surface roughness Ti\u20136Al\u20134V specimens. A temperature gradient exists between the pit and ridge area for each specimen, which is the highest in the high-roughness specimen as shown in Fig. 15(d)\u2013(f). The heat flux for each specimen was also analyzed, and the maximum heat flux was compared as shown in Fig. 15(g)\u2013(l). The highest maximum heat flux is shown in Fig. 15(i) and l, which is due to the high-roughness specimen. These mechanical and thermal computational analyses show that the surface roughness affects the tensile stress distribution and the thermal heat flux distribution under the same boundary condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000089_ev.2019.8892977-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000089_ev.2019.8892977-Figure5-1.png",
        "caption": "Fig. 5. PMVM-SMM magnetic flux density map and flux lines distribution",
        "texts": [
            " a) PMVM with surface magnets mounted in rotor and stator poles (PMVM-SMMRS) The outer region is based on 27 negative and 27 positive magnets oriented radial which gives an advantage in magnetizing the magnetic core. Due to the both stators, the rotor is placed between the stators and receives magnetic energy from two sources to interact with PM stuck inside. The core material used for all structures is M270 with 0.35 thickness and the PM are 1.4 T. In this section it is presented the magnetization of the cores and the flux lines passing through inner stator, the inner air gap, the PM in the rotor, the rotor core, the outer air gap, the PMs in the outer stator and the outer stator core. From Figure 3 to Figure 5, a) can be highlighted that the magnetic flux density does not exceed 2.3 T, although PMVM-SMMRS rotor and outer stator is more saturated in comparison with PMVM-BM and PMVM-SMM. The flux lines distribution is represented in Figure 3 to Figure 5, b) where the role of PM can be highlighted. Not all the flux lines produced in the stator secondary poles are closing through the outer stator which results in the influence of the coercive magnetic field produced by the PM in the outer stator. Following the electromagnetic analysis, electromagnetic torque variation in time is represented in Figure 6. PMVMSMMRS develops 108 Nm, an average value of torque with a consistence of 13.35% torque ripples and it is rated at 1696.5 W output power with 95"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001861_1350650120972499-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001861_1350650120972499-Figure11-1.png",
        "caption": "Figure 11. The cloud chart of friction coefficient and load carrying capacity before optimum design.",
        "texts": [
            " The detailed derivation process of force equilibrium equations are shown in Tang et al.24 In this study, the force equilibrium equations of the slipper are written as follows Fp \u00fe Fs \u00fe FTP\u00f0 \u00de cosb\u00fe Fz h; _h; t \u00bc 0 Mx h; _h; t \u00feM1 h; _h; t \u00bc 0 My h; _h; t \u00feM2 h; _h; t \u00bc 0 8>< >: (33) Thus, the load carrying force and friction coefficient before optimum design can be obtained by force and torque module. The results of initial load carrying force and friction coefficient are set to input parameters in multi-objective optimization module. Figure 11 shows the cloud chart of friction coefficient and load carrying capacity before optimum design. In Figure 11(a) and (b), according to the range of texture parameters given above, the optimal design of texture is also to find the optimal solution of texture characteristics within this range. The optimal texture feature factors, i.e. the coordinates of the corresponding feature space, are found in the original constraint feature space. In this article, spherical dimple texture effects on load carrying and friction forces of slipper bearing were analyzed by simulation and experiment. Figure 12 shows textured slipper bearing test equipment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002757_s11786-021-00511-6-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002757_s11786-021-00511-6-Figure1-1.png",
        "caption": "Fig. 1 Basic coordinate systems",
        "texts": [
            " In mechanics, computer algebra is widely used to study the equilibria of differential and dynamical systems. Some computer algebra algorithms for solving these problems were described in [15]. The study of regions in the parameter space with certain equilibria properties also occurred in relevance to a biology problem was presented in [16]. We consider the system of two bodies connected by a spherical hinge that moves in a circular orbit [7]. To write the equations of motion of two bodies, we introduce the following right-handed Cartesian coordinate systems (Fig. 1): OXY Z is the orbital coordinate system, the OZ axis is directed along the radius vector connecting the Earth center of massC and the center of mass O of the two\u2013body system, the OX axis is directed along the linear velocity vector of the center of mass O , and the OY axis coincides with the normal to the orbital plane. The axes of coordinate systems O1x1y1z1 and O2x2y2z2, are directed along the principal central axes of inertia of the first and the second body, respectively (Fig. 1). The orientation of the coordinate system Oi xi yi zi with respect to the orbital coordinate system is determined by the aircraft angles \u03b1i (pitch), \u03b2i (yaw), and \u03b3i (roll) (see [7]) in the form a(i) 11 = cos\u03b1i cos\u03b2i , a(i) 12 = sin \u03b1i sin \u03b3i \u2212 cos\u03b1i sin \u03b2i cos \u03b3i , a(i) 13 = sin \u03b1i cos \u03b3i + cos\u03b1i sin \u03b2i sin \u03b3i , a(i) 21 = sin \u03b2i , a(i) 22 = cos\u03b2i cos \u03b3i , (1) a(i) 23 = \u2212 cos\u03b2i sin \u03b3i , a(i) 31 = \u2212 sin \u03b1i cos\u03b2i , a(i) 32 = cos\u03b1i sin \u03b3i + sin \u03b1i sin \u03b2i cos \u03b3i , a(i) 33 = cos\u03b1i cos \u03b3i \u2212 sin \u03b1i sin \u03b2i sin \u03b3i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003191_itec51675.2021.9490116-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003191_itec51675.2021.9490116-Figure5-1.png",
        "caption": "Fig. 5. Flux density distribution of the proposed five-phase synchronous reluctance motor. (a) peak power operation. (b) continuous power operation.",
        "texts": [
            " Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply. the proposed five-phase synchronous reluctance motor with harmonic current injection. The key parameters are summarized in TABLE 2. Based on the analysis in the previous section, a simple salient pole reluctance rotor is more suitable for the proposed five-phase SynRM with harmonic current injection. Fig. 4 shows flux distribution of proposed five-phase SynRM with simple salient pole reluctance rotor exited by d-axis and q-axis current respectively. Fig. 5 shows flux density under peak power and continuous power operation, respectively. Authorized licensed use limited to: Hoseo Univ. Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply. Fig. 6 shows the exciting current waveform of symmetric five-phase windings. Fig. 7 shows the phase current waveform with a different ratio of third harmonic current. It should be pointed out that the current (rms value) of third harmonic and fundamental components keep unchanged to maintain the same copper losses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001859_j.promfg.2020.10.106-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001859_j.promfg.2020.10.106-Figure5-1.png",
        "caption": "Fig. 5. Doctor blade subassembly model.",
        "texts": [
            " It is proven that this has a positive effect on mechanical properties of the finished part, such as density and tensile strength [14]. There are also systems which feature only a doctor blade for powder spreading [10]. The blades come in various geometrical shapes with particular advantages and disadvantages. In order to experiment appropriately, it is necessary to be able to test doctor blades of different geometries and materials and at different inclinations. The suggested PDS subsystem can be seen in Figure 5. Thus, it is possible by design to adjust the height of the doctor blade via the nuts that secure the z-spacers onto the base connectors. This controls the powder layer thickness that the doctor blade creates. The doctor blade is secured in a perfectly horizontal position by the special slot, which is bolted on the doctor blade rotator rod. The latter can be turned and secured at the desired angle of inclination by the tightening screws of the z-spacers. Moreover, the apparatus allows for a small inclination of the blade around z-axis on the horizontal plane",
            " It is proven that this has a positive effect on mechanical properties of the finished part, such as density and tensile strength [14]. There are also systems which feature only a doctor blade for powder spreading [10]. The blades come in various geometrical shapes with particular advantages and disadvantages. In order to experiment appropriately, it is necessary to be able to test doctor blades of different geometries and materials and at different inclinations. The suggested PDS subsystem can be seen in Figure 5. Fig. 5. Doctor blade subassembly model. Thus, it is possible by design to adjust the height of the doctor blade via the nuts that secure the z-spacers onto the base connectors. This controls the powder layer thickness that the doctor blade creates. The doctor blade is secured in a perfectly horizontal position by the special slot, which is bolted on the doctor blade rotator rod. The latter can be turned and secured at the desired angle of inclination by the tightening screws of the z-spacers. Moreover, the apparatus allows for a small inclination of the blade around z-axis on the horizontal plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000790_10400435.2020.1744771-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000790_10400435.2020.1744771-Figure1-1.png",
        "caption": "Figure 1. Modifications made to the standard four-wheeled walker.",
        "texts": [
            " Acc ep ted M an us cri pt A custom experiment walker (EW) was developed by modifying a standard fourwheeled walker (Symphony, Shima Seisakusyo, Japan). The original four-wheeled walker weighed 6.4 kg and its width and length measured at 52.5 cm and 55.0 cm. Its handle height was adjustable, with a height range of 77 cm to 87 cm. The main modifications made were the removal of its braking system and the replacement of its rear wheels with a drive wheel assembly. All other additional components were placed in the basket component of the four-wheeled walker (Figure 1). As shown in Figure 1, each drive wheel assembly contains an iron core polyurethane drive wheel (UB100, Chubo Sangyo, Nagoya, Japan) connected to a shaft supported by ball bearings whose housing is attached to a custom-made aluminum base plate. This base plate was clamped onto the leg of the four-wheeled walker and was supported by a steel dowel pin pressed into the existing axle hole in the leg. The wheel is driven by a motor supported by the base plate an the shaft of the drive wheel was connected to the motor shaft using a flexible coupling (MJT-40CK-BL-12-18, Nabeya Bi-tech Kaisha, Japan)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000229_042014-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000229_042014-Figure1-1.png",
        "caption": "Figure 1. The spacecraft with the movable module.",
        "texts": [
            " The main aim of this work is to develop the simplest control algorithm, which can be applied in cases of simple constructional forms of the small spacecraft with primitive equipment, including nano-satellites, that is economically profitable and can be used in more space missions. Research of the attitude dynamics of the spacecraft still is the important task of the modern space flight dynamics, which topicality increases due to developing the modern schemes of the nanosatellites with simplest control systems, including the control systems with the movable internal masses [1-5]. In this paper the attitude dynamics spacecraft (SC) with a movable module on retractable beams of variable length (figure 1) is considered [1, 2]. The attitude control is fulfilling by the length of the beams changing that tilts the movable module and translates the position of the mass center and creates the torque from the jet-engine with constant fixation relative the main body of the spacecraft. The movable module can represent the multifunctional element, e.g. a telescope, radiometer, etc. Such a control scheme allows using the functional a module as the simple actuator of the control system. Therefore, in this case the control does not need the presence of special executive bodies. The main aim of this research is to develop the control scheme to suppressing the nutational oscillations with the help of impulses of jet-engine at the the constancy of the angular displacement of the movable module relatively the main part of the SC. Let us consider the compound structure of the SC [1, 2] with one movable module (Figure 1), which is connected by the retractable beams with variable length (1 - the main body of SC, 2 - the movable module, 3 - the jet-engine, 4 - retractable beams). The sequence of rotation angles of the main body of the SC relative the inertial space is following: Z X Z  , by angles  , , . The coordinates systems are: ITNT 2019 Journal of Physics: Conference Series 1368 (2019) 042014 IOP Publishing doi:10.1088/1742-6596/1368/4/042014 At small angles of the rotation of the movable module it is possible to neglect the terms from the curvature of retractable beams"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001197_s0263574720000533-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001197_s0263574720000533-Figure3-1.png",
        "caption": "Fig. 3. The backbone curve of the base part and the support polygon constructed by the base part.",
        "texts": [
            "1017/S0263574720000533 Downloaded from https://www.cambridge.org/core. University of Exeter, on 03 Jul 2020 at 02:30:12, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. polygon, the circle has a largest coverage area in all the plane polygons. So we choose to design the backbone of the base part as a planar arc. Then we use the backbone curve theory to discretize the backbone curve to generate the joint angles of the base part. The support polygon and the backbone curve are shown in Fig. 3. For the planar arc, the arc length, which equals the length of the base part of the snake robot, is assumed as l, and the radius of the arc is r. The central angle of the plane arc is \u03b8 . The equations are r = l \u03b8 (6) S = 1 2 r2(\u03b8 \u2212 sin(\u03b8)) \u03b8 \u2208 (0, 2\u03c0) (7) where S indicates the area enclosed by the backbone curve of the base part, which is shown in Fig. 3. Then the area of the support polygon has a maximum value. When \u03b8 equals \u03c0 , S has the maximum value, 1/2\u03c0r2. So the central angle \u03b8 of the arc is determined to be \u03c0 and the radius r is calculated as l/\u03c0 . Then the backbone curve of the base part of the snake robot can be discretized to calculate the joint angles for the base part. The discretization method based on the curvature and torsion of the backbone curve is given as,8, 20, 28 { \u03bay(s) = \u03ba(s) cos(\u03c6(s)) \u03bap(s) = \u03ba(s) sin(\u03c6(s)) (8) \u03c6(s) = \u03c60 + \u222b s 0 \u03c4(s)ds (9) where s is the arc length of the backbone curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002821_vlsi-dat52063.2021.9427332-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002821_vlsi-dat52063.2021.9427332-Figure4-1.png",
        "caption": "Fig. 4 The common bearing fault categories: (a) Healthy, (b) Inner ring defect, (c) Inner ring break, (d) Outer ring defect, (e) Outer ring break, (f) Cage break, and (g) Ball break.",
        "texts": [
            ", ADLINK USB-2405 is used in this experiment). On the contrary, the proposed method gathers the vibration signal through the sensor (i.e., ADcmXL3021 in this experiment) on the PYNQ-Z2 FPGA. At last, the screens are used to display the results of traditional and proposed methods, respectively. Based on the structure of the bearing, the target bearing fault categories are 1) healthy, 2) inner ring defect, 3) inner ring break, 4) outer ring defect, 5) outer ring break, 6) cage break, and 7) ball break, as shown in Fig. 4. With the identical bearing specification in Table 1, the goal of the following experiments is to detect the different bearing faults in Fig. 4. The frequency of the rotating speed of the motor is configured to 1,000 rpm as a design example. According to the number of data, we collect 360 data during each kind of bearing operation. Therefore, we have 2,520 data for this experiment, and each data include 10,240 points to represent the vibration signal. Furthermore, we randomly select 75% of the data (i.e., 1,890 data) as the training dataset and the remaining 25% data (i.e., 630 data) are used as the testing dataset. B. Verification result of the proposed embedded bearing fault detection platform As mentioned before, the PS unit is used to preprocess the receiving vibration signal, which is a time-domain signal, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure29-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure29-1.png",
        "caption": "Fig. 29. Rotor positions for measuring UMF. (a) STPSPM machine lacking 2 poles. (b) STPSPM machine lacking 1 pole.",
        "texts": [
            " A screw is employed to adjust vertical position of the rotor and a laser distance sensor is used to measure the vertical position of the rotor. Based on such test rig, the UMF in the vertical direction under different eccentricity ratio can be measured. To simplify the experiments, the UMF for both STPSPM machines is measured under the open-circuit condition. The weight of rotor, rods and movable platforms is measured beforehand and compensated in the measured UMF results. In the FE simulation, the rotor positions are adjusted deliberately in the same direction, which are shown in Fig. 29. As shown in Fig. 30(a), when the rotor is at the centric position, the UMF is negligible for the STPSPM machine lacking two poles. In addition, the measured UMF in the vertical direction increases with eccentricity ratio, which matches well with the FE analysis. The difference is mainly caused by the frictions or the displacement error of rotor vertical position. For the STPSPM machine lacking one pole, even though the eccentricity does not happen, the UMF in the vertical direction is about -309N, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000631_iscmi47871.2019.9004295-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000631_iscmi47871.2019.9004295-Figure1-1.png",
        "caption": "Figure 1. Simplified rotary actuator sketch.",
        "texts": [
            " The general sequence for moulding with silicone rubber are: (i) Pattern or mould making; (ii) Mould preparation; (iii) Mould filling with liquid silicone rubber; (iv) Curing; and (v) Removal of cured part. Soft rotary actuators change in angle when inflated. Because their fabrication demands casting the actuator body as a single piece, their entire body were casted with soft silicone while the layers separating each air channel was made inextensible by embedding with paper inside (Figure 2). The geometry variables for the rotary actuators (Figure 1) are given by: big radius, R1, Small radius, R2, offset distance from center of circle, d, height of air channel, h, thickness of wall, t. The geometric equations describing the area, surface area and volume are described by the following equations: ( ) ( ) ( ) ( ) ( ) ( ) ( ) The actuator consists of several compartments arranged around the circumference of a circle that can be actuated using both positive and negative pressure. Table I shows the geometric values of the rotary actuators produced in the work"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000220_ecce.2019.8913300-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000220_ecce.2019.8913300-Figure5-1.png",
        "caption": "Fig. 5. Magnetic field distribution of the proposed modular TC-LPMVM. (a) Flux density distribution at 0 electrical degree. (b) Flux density distribution at 90 electrical degrees. (c) Flux line distribution at 0 electrical degree. (d) Flux line distribution at 90 electrical degrees.",
        "texts": [
            " degree 2/3 3/4 5/6 Through the analysis, it shows that there is a range for the arc-pole ratio to make the proposed TC-LPMVM have a considerable performance. After the optimization, the thrust ripple of the 3-phase motor is less than 5%, while thrust ripple of the 5- and 9-phase motor can be less than 1%. Main design parameters of the optimized proposed modular TC-LPMVM are listed in the Table I. Andflux density and flux line distribution in the iron core and PMs under the listed condition are shown in Fig. 5. Through 3D FEM analysis, proposed one-phase module has a 14.9V no-load EMF, which is shown in Fig. 6(a). When loaded rated current, it will consume 175W electromagnetic power at average and output an average thrust force of 108.2N. The ripple of thrust force and electromagnetic power are reduced to a very low level, and will decrease as the phase number increases. Assemble 5 modules into a 5-phase motor and the distance between phases is \u0394=6/5 pole-distance. Thrust force and electromagnetic power of each phase and total machine of a 5-phase motor is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure1-1.png",
        "caption": "Fig. 1 The 2-DOF metamorphic mechanism and its composition principle",
        "texts": [
            " According to structural theory and formation methodology of metamorphic mechanisms based on augmented Assur groups (Li and Dai 2010), the constrained metamorphic mechanisms can be divided into four parts including 1 3 frame, active parts, several basic Assur groups and at least one augmented Assur groups. To illustrate the Assur group based metamorphic mechanisms, the 2-DOF metamorphic mechanism is taken as an example. The 2-DOF metamorphic mechanism is composed of active part, augmented Assur group RRRR and Assur group RRP (Fig.\u00a01b, d) and its composition principle are shown in Fig.\u00a01. The mechanism has two loops. The active part and frame are connected with augmented Assur group RRRR to form a loop. The other loop is composed of Assur group RRP connected to the former loop and frame. The metamorphic process of mechanism based on augmented Assur group is the process of transforming 1-DOF augmented Assur group into basic Assur group. The 2-DOF metamorphic mechanism shown in Fig.\u00a01 consists of an augmented Assur group, and its four metamorphic configurations are shown in Fig.\u00a02. According to the above-mentioned structural theory and formation methodology of metamorphic mechanisms based on augmented Assur groups, the unified dynamics modeling method of constrained metamorphic mechanism is studied. The driving forms of active parts are shown in Fig.\u00a03. Figure\u00a03a and b shows active parts in pure rotational and pure prismatic form, respectively. According to the Newton\u2013Euler equation (referred to as N/E equation), the dynamic equations of active parts can be written as follows, where Fi\u22121,i and Fi+1,i are the force vectors at component Li exerted by components Li\u22121 and Li+1 , respectively, and (1) { Fi\u22121,i \u2212 Fi,i+1 + Fi = miC\u0308i Mi\u22121,i \u2212 li \u00d7 Fi\u22121,i \u2212Mi,i+1 \u2212 hi \u00d7 Fi,i+1 +Mi = IC,i\ud835\udf3ai Fi+1,i = \u2212Fi,i+1 ; Mi\u22121,i and Mi+1,i are the moments at component Li exerted by components Li\u22121 and Li+1 , respectively, and Mi+1,i = \u2212Mi,i+1 ; Fi is the external force vector acting on the component Li ; Mi is the external moment acting on the component Li ; IC,i is the moment of inertia of the component Li around the centroid Ci ; i is the angular acceleration of the component Li ; C\u0308i is the acceleration vector at the centroid of the component Li"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002511_012004-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002511_012004-Figure7-1.png",
        "caption": "Figure 7. Meshed 2D Model of Spur gear. Figure 8. Model with 60% material removal.",
        "texts": [
            " Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10.1088/1757-899X/1109/1/012004 This assembly 3D model (Figure 6) can be transferred or edited back and forth in any CAD/CAE software such as: CATIA, Inventor, and Solidworks, with similar dimensions. Gear topology optimization is performed with the finite element method using ANSYS software. After the driving gear has been designed in CAD software, the researcher will export to ANSYS software to analyse and optimize the design (Figure 7-10). We will proceed with keyhole cutting and disk creation in the Space Claim environment. A design calculation is performed with the model with 60% material removal (Figure 8). This is obtained after the material particles are removed and holes of the corresponding hole or groove shape are cut, as long as the fabrication conditions are fulfilled and aesthetic factors are guaranteed. It is then possible to replace the holes with corresponding large and small round holes, which is very convenient in machining, and optimal shape (Figure 9) is ensured"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002604_j.apacoust.2021.108063-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002604_j.apacoust.2021.108063-Figure4-1.png",
        "caption": "Fig. 4. FE model of door seal strip.",
        "texts": [
            " 3 reveals a comparison of the experimental complex modulus, the simulated complex modulus with and without considering the elastoplastic model. The simulated complex modulus considering the elastoplastic model agrees well with the experimental results under both 1 mm and 0.3 mm excitation amplitudes. However, without considering the elastoplastic model, the complex modulus is insensitive to the excitation amplitude. The above results demonstrate the validity of the elastoplastic model. The identified parameters are listed in Table 1. The typical cross-section of the door seal strip considered in the present study is illustrated in Fig. 4. The ring shape part is almost 11 mm in height with a thickness of 1.5 mm, and it is compressed when the door is locked on the vehicle body. The lower part is used to fix the seal strip to the vehicle body, with a U-shaped metal bone strengthening the whole component. Considering the practical loading conditions, the seal strip can be modelled as a twodimensional, plane strain FE model, as shown in Fig. 4. The element size is specified as 0.5 mm so that a satisfactory mesh refinement is achieved, and at the same time the computation cost is acceptable. In this figure, the black line represents the door, which is considered infinitely rigid during the door closing process. The lower part is clamped to the ground, which represents the vehicle body. For saving the computation time, the lower part only meshed where contact with the upper ring is expected during the compression process. Thus, the door with the sheet metal is modelled through an analytical rigid panel, and the seal is modelled with 712 nodes and 576 four-node plane strain elements with the hybrid formulation. As depicted in Fig. 4, the blue part is made of foam rubber. The green area is dense rubber, and the red region is steel material. In the actual compression process, it is nearly prohibited interpene- ial: (a) DMA test and (b) simulated method. Table 1 Identified parameters of overlay constitutive model. C10 (MPa) D1 gi si Ep (MPa) rpy (MPa) 0.1328 0.0001 0.8191 0.0021 0.9850 0.0071 tration between the rubber and sheet metal surface, and the rubber itself. In this numerical simulation, there are 3 contact pairs, and penalty-based slipping friction [25] is used with friction coefficient 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000963_s00354-020-00093-0-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000963_s00354-020-00093-0-Figure5-1.png",
        "caption": "Fig. 5 Multi-agent system",
        "texts": [
            " Finally, chemotaxis-inspired formation control can be found in [4, 10]. However, their methods cannot directly control the relative positions of agents and, therefore, cannot be applied to our problem. Formation Control Inspired by\u00a0E. coli Chemotaxis This section presents a chemotaxis-inspired solution to the formation control problem of multi-agent systems. In particular, we focus here on the chemotaxis controller (2) of E. coli and present a formation control method based on it. Consider the multi-agent system illustrated in Fig.\u00a05, which consists of n agents. Agent i ( i \u2208 {1, 2,\u2026 , n} ) is represented by the model of a two-wheeled mobile robot shown in (1), that is, where xi(t) \u2208 \u211d 2 is the position, i(t) \u2208 \u211d is the orientation, and ui1(t), ui2(t) \u2208 \u211d are the control inputs determining the translational and rotational velocities, respectively. We introduce a local controller Li to each agent i. This is of the form where i(t) \u2208 \u211d m is the state, [xj(t) \u2212 xi(t)]j\u2208\u2115i \u2208 \u211d 2|\u2115i| is the input, ui(t) \u2208 \u211d 2 is the output, i.e., ui(t) \u2236= [ui1(t) ui2(t)] \u22a4 , and gi1 \u2236 \u211d m \u00d7\u211d 2|\u2115i| \u2192 \u211d m and (5) \ufffd xi(t + 1) i(t + 1) \ufffd = \ufffd xi(t) i(t) \ufffd + \u23a1\u23a2\u23a2\u23a3 cos( i(t) + ui2(t))ui1(t) sin( i(t) + ui2(t))ui1(t) ui2(t) \u23a4\u23a5\u23a5\u23a6 , (6)Li \u2236 { i(t + 1) = gi1( i(t), [xj(t) \u2212 xi(t)]j\u2208\u2115i ), ui(t) = gi2( i(t), [xj(t) \u2212 xi(t)]j\u2208\u2115i ), gi2 \u2236 \u211d m \u00d7\u211d 2|\u2115i| \u2192 \u211d 2 are functions specifying the structure of the controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001354_s10015-020-00628-0-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001354_s10015-020-00628-0-Figure2-1.png",
        "caption": "Fig. 2 The specifications of the robot",
        "texts": [
            " There are always five food objects and five poison objects in the field. A new food or poison object will be generated with a random position when it is transported to the nest. The radii of food and poison objects are set to 5.0\u00a0m and 2.5\u00a0m, respectively. The food and poison objects are set to have the same weight; in particular, at least four robots are required to move an object. Since a robot does not have the ability to distinguish between the two objects, the only difference between them are in the size. The specifications of the robot are shown in Fig.\u00a02. Each robot is composed of eight distance sensors, an omnidirectional camera, an artificial neural network controller, and two motors to rotate the left and right wheels. The range of the distance sensor is set to 3.0\u00a0m, detecting the distance between the nearest objects, robots, or walls. The range of the omnidirectional camera is set to 15\u00a0m, gathering one input for the distance and two inputs for the relative angle, represented by the sine and cosine of the angle, of the followings: \u2013 The nearest food or poison object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000503_012020-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000503_012020-Figure5-1.png",
        "caption": "Figure 5. Flow velocity after passing the water wheel",
        "texts": [
            " The red color shows the highest flow velocity area on the waterwheel blade 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 with a maximum value of 3.42 m/s, while the lowest water flow velocity that affects the waterwheel blade is light blue with a value of 2.158 m/s. The results of numerical analysis show that from the simulation results of the waterwheel obtained at 2.19 m/s, the speed of water flow after passing the water wheel shown in Figure 5 is 5.076m/s, and the maximum pressure experienced by each blade of water is 42326 Pa. The phenomenon of water flow in the waterwheel has been studied using CFD simulation using Ansys Student Version 19.1 calculated based on flow assumptions, with an initial flow velocity of 5.00 m/s by entering the gravity acceleration value 9.81 m/s2. The position of the water wheel is assumed to be half float. The model of approach used is k-epsilon. By using water-liquid fluid material, stating that the waterwheel moves based on the flow of water by 3rd NICTE IOP Conf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000234_smc.2019.8913932-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000234_smc.2019.8913932-Figure8-1.png",
        "caption": "Fig. 8: 3D view of Berrick\u2019s eye. In order: support structure, neopixel lighting system, camera, eye cover.",
        "texts": [
            "4V, thus we use a DCDC Step Down 10A component to reduce the voltage to 6V with an efficiency of 95% (3W consumption at full load when providing 57W). D. Vision system One of the main contributions of this project is the exploration of an innovative HRI gazing system, based on eyes lighting to provide feedback to the human on the accomplishment of the vision task. The mechanical structure of the eyes has been completely redesigned to accommodate the Neopixel lighting source. The 3D model of the eye\u2019s structure is shown in Fig. 8. Motion of the eyes is achieved by combining two servomotors: one is directly connected with the ocular system and is responsible for tilting, the second is connected with both eyes individually with a cinematic structure that ensures consistent motion of the pair, i.e. both eyes have always the same pan angle. The vision sensor is a Pi camera v1.3 able to deliver a 5MPx resolution image, or 1080p HD video recording at 30fps thanks to an Omnivision 5647 sensor and a fixed focus (1m-\u221e) lens of 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001555_physreve.102.032701-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001555_physreve.102.032701-Figure5-1.png",
        "caption": "FIG. 5. Schematic diagram of a section of stripe above the SmC to SmA transition temperature TAC. The shaded areas of thickness \u03b4 at the top and bottom surfaces are assumed to have a nonzero tilt order, while the rest of the medium is in the SmA phase without any tilt order. As described in the text, in order to reduce the positive elastic energy of bending of the layers, the center OA is pushed far away from the surface, increasing Ri(= OAL) to tens of micrometers.",
        "texts": [
            " 3, a change in the number of stripes would require a major structural rearrangement of the bent layers, which is a slow process. As the sample is heated above the SmC to SmA transition temperature TAC, the tilt order is lost in the bulk, i.e., in the interior layers of the FSF. However, the tilt order persists near the surfaces [14], decreasing in strength as the temperature is increased. We make the simplifying assumption that a relatively small and uniform tilt order persists over some thickness \u03b4 near the surfaces of the sample, and the order vanishes in the rest of the sample (Fig. 5). It is clear that, unlike in the SmC phase, the elastic terms favoring the formation of stripes which require the c-vector field are effective only in the two surface regions, whose relative contribution decreases as the thickness of the meniscus h increases. Consequently, in the free energy density averaged over the volume of a V-shaped section, while the kB term is multiplied by Ln((Ri + h)/Ri ) as in Eq. (10), the kKC and kC terms are multiplied by Ln[(Ri + h)(Ri + \u03b4)/Ri(Ri + h \u2212 \u03b4)], reflecting the reduced 032701-6 contributions from the latter terms",
            " The tilt order in the surface layers can be expected to be lower than in the SmC phase, and we use the following parameters in the calculations: RM = 200 \u03bcm, kB = 10 \u00d7 10\u221212 N, kSS = \u22123 \u00d7 10\u221212 N, kC = 0.6 \u00d7 10\u221212 N, kKC = 4 \u00d7 10\u221212 N, and \u03b4 = 20 nm. Assuming that the stripe domain emerges from the two-dimensional structure at h = 25 \u03bcm, the energy density in the absence of the stripes, arising from the kSS contribution [Eq. (1)] is 0.0012 J/m3. The dominant layer bending contribution can be reduced by increasing Ri (Fig. 5). As the stripe structure is found to smoothly emerge from the focal conic structure, with the periodicities in the two being comparable [15], the angle \u03b1 has to be reduced as Ri is increased (Fig. 5). The average energy density decreases as Ri is increased, the rate of change decreasing as well. When Ri = 80 \u03bcm, and \u03b1 = 0.18 rad, the energy density is just lower than that of the meniscus without the stripe. This can be considered as the threshold condition for a stable stripe structure. For a smooth structure in the meniscus, as in the SmC phase, we can assume that Ri and \u03b1 remain independent of h. As Ri is much larger than h, it is clear from Eq. (13) that the stripe width w is not very sensitive to h, unlike in the SmC phase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000162_012003-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000162_012003-Figure1-1.png",
        "caption": "Figure 1. Real (a) and modelled (b) suspension systems.",
        "texts": [
            " Initially, the current material, loading conditions and constraints to which the control arm is subjected were determined, taking into account its traditional design. Secondly, the analysis of the part was developed considering its current material and shape to obtain the total displacement and von Mises stress distributions and their maximum values. Then, the new material model was defined, and the shape optimization was performed with the goal of maximizing stiffness. Next, the results from the optimization process were validated by manufacturing and testing the printed part. Finally, conclusions were formulated. The suspension system of a buggy car, Figure 1(a), was taken as a reference for the study. This system was analysed to obtain the forces on the lower control arm, for the calculation of stress and displacement fields with its current design using Ansys v18.1. The geometry used to study the current model can be seen in Figure 1(b). The material of the control arm was steel, one of the most used to build this mechanical part in commercial automobiles. FEA requires the appropriate definition of boundary conditions and mesh for optimal solution [15]. The Neumann and Dirichlet boundary conditions corresponded to the loads applied to the control arm and its supports, respectively. To obtain the reactions forces in the component that was analysed, a kinetic analysis was performed using SolidWorks Motion complement. In this process, the following parts of the system were modelled: the upper control arm, the lower control arm, the spring shock absorber, and the ball joints. A quarter-vehicle model was taken into account to perform the study since it is a simple model but widely used in suspension design [16]. The assembly shown in Figure 1(b) was used to calculate the reactions in the lower control arm. In this case, the spring shock absorber is not connected to the lower arm as usual, but to the upper one. An upward vertical force of 1125 N was applied in the middle of the steering knuckle, this magnitude corresponds to a quarter of the weight of the vehicle (including passengers). For the simulation, the spring stiffness used was 30 N/mm and the damping coefficient of 5th IMRMPT Journal of Physics: Conference Series 1386 (2019) 012003 IOP Publishing doi:10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002215_s41693-020-00051-8-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002215_s41693-020-00051-8-Figure5-1.png",
        "caption": "Fig. 5 Adaptive Pick and Place setup, with an ABB 1600 robot and brick conveyor system. The robot picks up the brick from the corresponding conveyor track and passes it through the fork sensor, before positioning it on the panel substrate on which mortar is applied. The panel is fixed to the table using a CNC milled jig",
        "texts": [
            " Once the bricks have been formed in the factory, they are packaged and transported to the jobsite where they are to be pre-assembled onto panels. Although robotic pick and place is a highly mature technology which is commonly used in manufacturing environments, there were various challenges we faced in adapting this technology to the project. First, we are using 14 types of bricks and they are not stackable, so while standard pick and place operates on a single conveyor, we developed a cassette system which houses our different families (Fig.\u00a05). The conveyor system is passive and fitted with roller tracks. The tracks are tilted to ensure the bricks are presented to the robot in the calibrated pickup point. The panels feature 16 bricks each, usually consisting of 3\u20134 different brick families which are fed correspondingly into the different tracks. The robot operator has an assembly sheet information and refills the bricks as needed. This rack system could be improved in further implementations. The second difficulty we encountered was the problem that the curving geometry of the bricks did not allow for the usage of an industry-standard vacuum gripper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000297_icems.2019.8922215-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000297_icems.2019.8922215-Figure2-1.png",
        "caption": "Fig. 2. The feasible fabrication process of VLW LDMOS. (a) etching trench, (b) depositing high-k dielectric. (c) Forming P-well, field region, poly and S/D region. (d) metalization. (e) fabrication flow of high-k dielectric region.",
        "texts": [
            " (1) reduces to: Wd Cbox Ecx \u2248 q Nd ts Ws (2) Equation (2) provides a useful approach to optimize the device parameters of SOI LDMOS. Obviously, a linear Nd , ts or Ws can ensure the equality of Eq. (2). Actually, the linear Nd , and ts have already been proven as effective methods to obtain the ideal lateral electric field and applied in the design and manufacture of SOI LDMOS [3]\u2013[5]. In this letter, a linear Ws is first proposed and a so-called VLW technique is developed to even the electric field in the drift region of SOI LDMOS. A feasible fabrication process of the VLW LDMOS is exhibited in Fig. 2. In this instance, PZT is used to form the high-k region due to its high enough permittivity and extensive study [15]. The process begins by etching trenches as shown in Fig. 2(a), which is compatible with the dielectric isolation process in SOI CMOS technology. Figure 2(b) shows the 3D view after PZT region is formed. The key steps to form the PZT region is shown in Fig. 2(e). Firstly, the dry oxidation is implemented to obtain a thin SiO2 buffer layers at the high-k/silicon interface [11]. Then, PZT is deposited into the trenches by pulsed-laser deposition (PLD) technique [15]. Following, the wafer surface is planarized by the chemical mechanical polishing (CMP) [16]. After this, as shown in Fig. 2(c), P-well region is formed by boron implantation. The gate stack is formed by the gate oxide and polysilicon deposition followed by patterning. Then, the N+ source/drain region and P+ region are formed by phosphorus and boron implantation, respectively. Finally, the metal electrodes are formed by the standard deposition, patterning and metallization techniques in CMOS process. Figure 2(d) illustrates the final 3D structure. Obviously, the long-time and high-temperature annealing is not needed in comparison with the VLD and VLT processes. A 3D semiconductor device simulator, Synopsys TCAD, is used to investigate the proposed novel device. The potential contours of the CONV and VLW LDMOS with the same geometry parameters (e.g. ts , tox , Wd and Ld) are shown in Fig. 3(a). In the CONV LDMOS, the potential lines crowd at the ends of the drift region, which leads to a high electric field and premature breakdown"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001869_s11223-020-00217-3-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001869_s11223-020-00217-3-Figure2-1.png",
        "caption": "Fig. 2. Static finite element model of the wheel.",
        "texts": [
            "1007/s11223-020-00217-3 published by Wang and Zhang [3], and it focuses on seeking a practical and effective method for predicting the wheel fatigue life in the test by using a special fatigue analysis software based on the local stress-strain approach. The conclusions of this study are made fully in accordance with the new methods and new findings described in detail. 1. Establishment and Calibration of a Static Finite Element Model of the Wheel. The studied wheel, produced by Dongfeng Automotive Wheel Co., Ltd., is a passenger car wheel made of steel. Figure 2 depicts shows a static finite element model, with plate and shell elements applied to the rim and disc model. The detailed information on dimensions of the rim and disc is given in Table 1. The material properties were as follows: Poisson\u2019s ratio 0.3, elastic modulus E 214 GPa, and density 7900 kg/m 3 . The material of this wheel is equivalent to St37 steel, whose fatigue properties are listed in Table 2 [8]. As shown in Figs. 1 and 2, the test rig was restrained, and a force of 1562.4 N was applied to the center of the moment arm with a length of 938 mm",
            "4 N with F0 being the amplitude of F ty ( ) or F tx ( ), nY y, is the normal Y strain of element Y obtained via static finite element analysis applying only PFEA x, , nX y, is the normal X strain of element X obtained via static finite element analysis applying only PFEA y, , PFEA x, is the static force corresponding to F tx ( ), P FFEA x, 0 2344.4 N, nY x, is the normal Y strain of element Y obtained via static finite element analysis applying only PFEA x, , and nX x, is the normal X strain of element X obtained via static finite element analysis applying only PFEA x, . finite element model of the wheel shown in Fig. 2. Thus, the calculation of strain histories is based on static finite element analysis in the fatigue analysis. In theory, the calculation accuracy can be high enough only when the loading frequencies are sufficiently lower than the first-order mode frequency of the structure. The earlier introduced transient finite element model of the wheel was used to study the effect of loading frequencies of F tx ( ) and F ty ( ) on the amplitude of the normal Y strain history of element Y and normal X strain history of element X "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001118_10426914.2020.1772489-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001118_10426914.2020.1772489-Figure7-1.png",
        "caption": "Figure 7. Spur gear with 20\u00b0-20\u00b0.",
        "texts": [],
        "surrounding_texts": [
            "Specimens prepared to study WCEDM process parameter and the gear blanks used to cut the gears are made up of with EN353 material. The material composition of procured raw material i.e. EN353 rod was analyzed using an optical emission spectroscope (OES). Three samples of size 20 mm diameter and 5 mm thick prepared from the raw material and polished the surface with 60-grit emery paper. The sample was placed on the spark stand and flooded with inert gas and serious of sparks have been created on the grounded surface of the sample. The elements of the raw material are calculated from the recorded signals in OES. The result of the OES elemental analysis of the procured raw material is given in Table 1. In room temperature the mechanical properties of EN353 steel at room temperature in annealed condition are given in Table 2. The quality of the product machined from the WCEDM process is always influenced by various machining parameters most importantly Ip, Ton, Toff and SV. These influence the performance parameters Ra and MRR. Hence, proper selection of machining parameters leads to higher MRR, lower Kf and lower Ra. In the present work, the process parameters such as Ip, Ton, Toff and SV were chosen each at three levels as given in Table 3. Taguchi\u2019s L9 orthogonal array used in this study is given in Table 4. A brass plain hard wire of 0.25 mm diameter connected to the negative polarity was used to machine the prepared samples connected to the positive polarity. A series of cuts were made in the samples using four axis computer numerical control controlled WCEDM machine (Maxicut-E, Electronica machine Tools Ltd, India). It consists of Figure 1. Sample work piece-1. a machine tool, power supply unit and dielectric unit. The Machine uses an iso-frequent pulse generator; with a maximum operating current of 30A. The work volume of the machine was 300 mm (X) x 400 mm (Y) x 225 mm (Z). The specifications of the WCEDM machine are given in Table 5. Clamps and bolts were used rigidly fix the sample on the work table. The IP, Ton, Toff and SV were chosen as input parameters based on the L9 array. The other influencing parameters such as work piece geometry, the thickness of the metal (10 mm), cutting length (10 mm), dielectric pressure, wire tension (780 g) and dielectric fluid were kept as constant throughout the experiments. The samples produced through WCEDM machine to carry out the present study are shown in Figs. 1 and 2. The performance characteristics such as Ra, Kf, and microstructures of the machined samples were investigated in this study. The Ra was measured by using a surface roughness tester (Mitutoyo, Surftest 301) has a 0.8 mm cut off length and a stylus radius of 0.0025 mm. Average of five readings were taken in the longitudinal and transverse directions are reported in this work. The machined surfaces are examined though an optical microscope in order to explore the effect of WEDM process on the micro structure of the material. The parameters which yield minimum Ra are used in the WCEDM to cut the gears used in this study. The involute profile of the gear from limiting radius to addendum radius for coast and load sides are drawn by using the following relation in AUTO CAD drafting software[5] From the geometrical details shown in Fig. 3, the following relations are obtained: The half tooth thickness at any radius ri is given by Pressure angle at any radii ri of the involute derived from the Eq. (2) is given as The angle subtended by the half tooth thickness at radius ri is given by The coordinates of the point i on the involute profile can be obtained by using Eqs. 6 and 7 by taking the gear center as origin and the tooth center line is collinear with y axis are Table 5. Wire cut EDM specification. Description Value Description Value Table traverse movement 300x400 mm Work piece weight (Max) 300 kg Aux. table traverse movement 80 x 80 mm Resolution 0.0005 mm Table Size 440 x 650 mm JOG speed (Max) 900 mm/min Taper angle (Max) \u00b130\u00b0/50 mm Wire spool capacity (Max) 6 kg (up to DIN 160/ p5) Work piece height (Max) 200 mm Wire diameter 0.25 mm (std.) 0.15,0.20 mm (opt.) Figure 2. Sample work piece-2. Figure 3. Involute tooth profile. Series of points are generated using the coordinates of the Eqs. 6 and 7 from the limiting circle radius (rl) to the addendum circle radius (ra) and the required involute curve is obtained by joining these points using spline curve. The same procedure is used to construct the profile of the coast side with various pressure angles. The base circles for the profiles and root circle are also drawn. A line is constructed from the involute profile limiting radius to a tangent point to the base circle. Two circles are drawn by considering line length as the radius for load side profile and host side profile of the adjacent tooth. A circle is constructed such a way that it is tangent to root circle and the two circles constructed as discussed above. The three circles and the tooth involutes profiles are shown in Fig. 4. The unwanted portions of the constructed lines are trimmed to get the complete gear tooth profile. The final profile is shown in Fig. 5. The gear parameters used in this study are given in Table 6. The AutoCAD DXF file is imported into the hyper mesh software. There the file is converted into neutral form (IGES). The IGES format of the profile has been imported to the machine to develop the WCEDM program. The gear blanks are case hardened and carburized with surface hardness value of 55HRC. Since the study of generated profile metrology is aimed at this work, only a few gear teeth were cut and inspected. For each case, two gears were prepared. The manufactured gears are shown in Figs. 6\u20138. The optimized WCEDM process parameters based on the previous study which gives less Ra was used to cut the gear tooth profile and the values of the process parameters used in this study are given in Table 7. The Ra of the machined gear tooth was measured along the face width direction using a surface roughness tester (Mitutoyo, Surftest 301) and the average Ra value obtained was 5.7883 \u00b1 1.003 \u03bcm. Lead, pitch and profile errors are very important parameters to define the quality class of the gear tooth produced through the WCEDM process and the errors were measured using the Gleason gear tooth inspection machine which is shown in Fig. 9. After the metrological inspection, the static load capacity of a gear tooth is experimentally estimated using a single tooth load test using a universal tensile testing machine. A test rig has been developed to load a tooth normal to the tooth profile. Few teeth were removed by using WCEDM process for easy asses for the loading arm to load the tooth. After fixing the gear firmly in a position such a way that the tooth is loaded normal to the loading point, the load was gradually applied with the loading rate of 1 N/sec. The load and the deflection of the gear tooth were monitored and plotted in a graph. Each gear type four sample teeth were loaded and the average result is reported in this work. The test set up is shown in Fig. 10"
        ]
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure15-1.png",
        "caption": "Fig. 15 Push-pulling the connector model with a long range and its modeling results",
        "texts": [
            "\u00a014, that for the second situation is shown in the upper-right, and that for the third situation is the same as the second one. Siemens NX failed to update the model for all of the three situations, while the proposed method can successfully update the model for all of them. The comparisons in case studies 1\u20135 are sufficient to show that the proposed method outperforms the state of the art in terms of robustness. Nevertheless, only translational push-pulls were used in these case studies, and the push\u2013pull ranges were small. One more case study (Fig.\u00a015) was thus carried out to show that this work also applies to rotational push\u2013pull with a long range in the edit. The case study shown in Fig.\u00a016 further demonstrates effectiveness of the proposed method by including multiple push-pulls in a row and different push\u2013pull types (translational and rotational). Specifically, four push\u2013pull edits were performed in a row, containing both the rotational push\u2013pull type (e.g., the first push\u2013pull operation) and the translational push\u2013pull type (e.g., the third push\u2013pull operation), as well as single-face push\u2013pull operations (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001303_acc45564.2020.9147239-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001303_acc45564.2020.9147239-Figure1-1.png",
        "caption": "Fig. 1. Schematic of a continuum robot composed of soft segments, along with its equivalent model as a serial configuration of multiple rigid-link RPR mechanisms. Local coordinate frames (Frenet-Serret frames [7]) are defined at the two ends of each soft segment.",
        "texts": [
            " In this method, the robot is divided into multiple segments that each deform in the shape of a circular arc, which is referred to as the Constant Curvature (CC) assumption. Although soft continuum robots generally can deform into shapes with variable curvature, many existing soft robots are designed to conform to the CC assumption [13], [14], which enables the derivation of closed-form kinematics and Jacobian formulations [7]. In the discrete method, each soft curved segment is modeled as an equivalent rigid-link RPR mechanism, as shown in Fig. 1. Several methods have been proposed for kinematic control of segmented continuum robots. The fitting algorithm was investigated in [15]. In [16], a modular control scheme was proposed to control the configuration of a segmented contin- 978-1-5386-8266-1/$31.00 \u00a92020 AACC 909 Authorized licensed use limited to: Cornell University Library. Downloaded on September 16,2020 at 09:04:41 UTC from IEEE Xplore. Restrictions apply. uum robot, aiming at reducing the computational load of the fitting algorithm by dividing the robot into multiple modules",
            " We analyze the stability and convergence properties of the controllers in Section IV for the general case of a robot with N segments. In Section V, we validate our position regulation and trajectory tracking controllers in simulation on 5-segment and 15-segment robots and illustrate the effect of the controller gain. The kinematics of the soft continuum robot are discussed in this section. The robot is composed of a set of segments connected to each other in a series configuration. The soft segments of the robot are replaced by equivalent rigid-link RPR mechanisms [31]. Figure 1 depicts a soft continuum robot with soft segments, which is equivalently modeled with multiple rigidlink RPR mechanisms connected in a series. To clarify the details of the model, Figs. 2 and 3 show the soft segmented robot and its equivalent rigid-link robot, respectively. Figure 2 illustrates a planar segmented soft continuum robot with N bending segments, each conforming to the constantcurvature (CC) assumption. The kinematic model of the continuum robot is defined by the kinematic equations of the multi-segment N -RPR rigid-link robot in Fig. 3. Since we assume that each segment of the robot is equipped with local sensors and actuators, it is able to measure the position of its prismatic joint and relative rotations of its revolute joints in its local coordinate frame. Furthermore, the i-th segment can communicate these local measurements to the adjacent (i\u2212 1)-th and (i+ 1)-th segments, as shown in Fig. 2. As shown in Fig. 1, the angular difference between the tangential local coordinate frames attached to the base and end-effector of the i-th segment and the orientation of the equivalent RPR rigid-link mechanism is defined as \u03b1i 2 . Furthermore, we denote the arc length of the i-th soft segment 910 Authorized licensed use limited to: Cornell University Library. Downloaded on September 16,2020 at 09:04:41 UTC from IEEE Xplore. Restrictions apply. by Li. Accordingly, as in [31], the position vector that is aligned with the prismatic joint of the i-th segment can be represented in the segment\u2019s local coordinate frame as the following vector ipi: ipi = [ Li sin(\u03b1i) \u03b1i Li 1\u2212 cos(\u03b1i) \u03b1i 0 ]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure31-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure31-1.png",
        "caption": "Fig. 31. A solid block with the same size as 6004 bearing.",
        "texts": [
            " The load sensor used in the measurement is Hangzhou MEACON MTK-LCS2, and the displacement sensor model is Keyence GT2-P12K. The load signal uses the analog channel of the NI USB-6002 DAQ card for acquisition, and the displacement uses the encoder channel of the MCC USBQUAD08 for acquisition. The test data is collected and stored by LABVIEW. The test bench itself also has deformation while measuring. The displacement of the fixture needs to be removed from the bearing displacement measurement data. To measure the displacement of the bench system under the measuring load, a solid block (shown in Fig. 31) is specially made with the same dimensions of the test bearing (6004 inner diameter, outer diameter, and width are 20 x 42 x 12 mm). The deformation of the solid block under the axial test load condition can be calculated using the FEM method, as shown in Fig. 32. The maximum displacement of a solid block under the load of 1000 N is 0.4 \u03bcm. The measured axial load-displacement curve of the solid block is shown in Fig. 33. According to Fig. 33, the axial displacement of the solid block under the axial load of \u00b11000 N is -3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure15-1.png",
        "caption": "Fig. 15 The general structure of the PHYSIOTHERABOT/w1.",
        "texts": [
            " EMG signals belonging to the patient are also recorded in the system. However, it is not used as a feedback element in the control cycle. At this point, the applications where the EMG participates in the control cycle are highly significant. The robot manipulator made of 7000 series aluminum material has three rotational axes. With these axes, flexion-extension, ulnar-radial deviation, and pronation-supination movements are performed. The general structure of the robot manipulator can be seen from Fig. 15. 430 Erhan Akdogan and Mehmet Emin Aktan The patient\u2019s arm is placed on the armrest. The patient\u2019s hand is placed between the bars in the handle. With the force sensor located just below the handle, the force and torque values that are applied by the patient are measured. With this sensor which can measure in six axes (three for force and three for torque), the force is measured during flexion-extension and ulnar-radial deviation movements, and the torque values are measured during pronation-supination movements. As shown in Fig. 15, there are mechanical limitations for safety in each axis. By means of the pins placed in these limitations, the ROM of the joints can be limited to the desired values. The kinematic and dynamic analysis of a robotic system is important for robot control. When performing kinematic and dynamic analysis of robots, classical manual calculation techniques can be used. However, as the DOF of the robot increases, these calculations become more complicated. For this reason, for three-DOF PHYSIOTHERABOT/w1, the analysis programs that can calculate the parameters related to the system model were used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure8-1.png",
        "caption": "Figure 8. Stress-strain state with cam axial position 0.11 mm less than when the shank is torn off",
        "texts": [
            " 7 it is visible that tensions in a rod in a contact zone with a package make from 300 MPa on edge of a head to 1000 MPa on a site of transition of a head to ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 a cylindrical part. In the sheet there are stresses of about 200 MPa. In the previous moment of calculation, differing on axial position of a cam on 0,11 mm, tensions in a rod in a contact zone with a package make from 100 MPa to 800 MPa, that is appreciably less, as shown on fig. 8. After detachment of the shank and removal of the pulling load from the rod, we observe the stress field shown in Fig. 9. The tension level in the contact zone of the rod with the package has changed. Now the stresses in the rod are from 100 to 400 MPa, and in the package - from 150 to 250 MPa. It should be noted that such a level of stress will occur when retracting the rivet bolt rod with a force as close as possible to the damaging force. If the retraction force of the rod is less, the stress in the contact area will be less, which will not increase the endurance of the connection as much as it is possible to do technologically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000470_ab606a-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000470_ab606a-Figure1-1.png",
        "caption": "Figure 1. Experimental setup consisting of a bar with equispaced suspension points and a smartphone. The labels are described in the text.",
        "texts": [
            " The period of the small oscillations, T, depends on the mass M, the distance from the suspension point to the center of mass R and the moment of inertia I as T = 2\u03c0 \u221a I MgR (1) where g is the gravitational acceleration. In the current study, small oscillations of a compound pendulum are analysed using different modern technologies [1\u20134]. The period of oscillation is measured as a function of its point of suspension, which affects its moment of inertia. As it involves key concepts in classical mechanics and can be readily implemented virtually in any Physics laboratory, the present experiment could encourage students\u2019 interest and motivation to experiment by themselves. The experimental setup, depicted in figure\u00a0 1, consists of a rigid metallic bar with equispaced holes and a smartphone. As there are several possible suspension points, the moment of inertia depends on the selected point. The dimensions of the bar, with holes made at points separated a uniform distance of 1.0 cm, are L = 1.199\u00a0 m and w = 0.024\u00a0m, and the mass M = 0.2518 kg. The smartphone is a Nexus 5, with mass m = 0.1311 kg, length Ls = 0.135\u00a0 m and ws = 0.068 cm. The distance from the suspension PED 1361-6552 10.1088/1361-6552/ab606a 2 Published 3 1361-6552/ 20 /023004+4$33"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003238_s12206-021-0829-0-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003238_s12206-021-0829-0-Figure1-1.png",
        "caption": "Fig. 1. Fan-driven gearbox.",
        "texts": [
            " Simultaneously, to accelerate the prediction capability of the temperature field in the preliminary design process of the gearbox, a numerical fitting method is used to establish the calculation formula for the surface convective heat transfer coefficients of each part in the planetary gear hole mode and compare them with the convective heat transfer coefficients in the non-opening mode, providing a design method for planetary gear hole parameters. A fan-driven gearbox is a herringbone star gear transmission. As shown in Fig. 1 [27], the input end of the gearbox is connected by splines, and the sun gear is a floating component. The planetary gear adopts a gear-bearing integrated design, and its internal support is a self-aligning roller bearing. The ring gear is a semi-floating member that is connected to the output shaft by bolts. The torque transmission bracket bears the gravity of the gearbox, and torque is generated when the gears mesh. To improve the oil-return capability of the fan-driven gearbox, the influence law of the planetary gear hole parameters on the heat transfer performance of the gearbox is determined by opening the planetary gear\u2019s undercut grooves and developing a design method for the planetary gear hole parameters, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002350_iros45743.2020.9341242-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002350_iros45743.2020.9341242-Figure1-1.png",
        "caption": "Fig. 1. Prototype of developed inverted-pendulum-type robotic wheelchair [12]. The left figure shows the wheelchair going up stairs and the right figure shows it going down stairs.",
        "texts": [
            " The iBOT climbs by alternating fourand two-point grounding, and provides a compact mechanism for going up and down stairs regardless of the stair slope and tread length. However, the statically unstable two-point grounding on stairs requires the user to hold the handrail to maintain balance while climbing up stairs without assistance of a caregiver [8]. Therefore, the iBOT is unsuitable for elderly people with limited lower- and upper-body function. The EPW proposed by Shino et al. [9] uses two-point grounding similar to the iBOT, as shown in Fig. 1, and climbs up stairs using the control theory of an inverted pendulum. This EPW has a slider mechanism to control the center-ofgravity position. Unlike the iBOT, it autonomously adjusts balance, omitting the necessity of caregiver assistance or user training. In a previous study using this stair-climbing EPW, a controller was implemented to maintain the pitch angle at 0 rad without falling on the stairs [9] [10] [11] [12]. Consequently, stability was guaranteed by taking approximately 5 s to actuate the rotary links for climbing upstairs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001535_j.jterra.2020.08.003-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001535_j.jterra.2020.08.003-Figure9-1.png",
        "caption": "Fig. 9. Multibody dynamic model of Tomcar TM27 vehicle (right) and side view of the vehicle passing over the obstacle (left).",
        "texts": [
            " The vehicle model is object-oriented, which makes it simple to incorporate changes such as number of axles, type of suspensions, and more. The second advantage is the ability to connect to external systems. The software is independent, so it can be integrated with any control software. In order to test the model, the results of simulation of an offroad vehicle, Tomcar TM27, using the new developed method, were compared with those obtained using the commercial Siemens\u2019 VL multibody software, which is based on the Lagrange equations. Fig. 9 shows the Tomcar TM27model that was developed in Siemens\u2019 VL. The points where the suspensions attach to the vehicle chassis were modeled using bushing (revolute, spherical, etc.). All other points of body-to-body connections were modeled using joints (revolute, spherical, etc.). Table 5 presents a list of the model joints and constraints and the bodies that were connected by each joint. Table 6 presents a list of the number of bodies and degrees of freedom of the four subsystems that comprise the model",
            " The mass of bodies, which are canceled, where added separately between the bodies that are connected to the canceled bodies. For the purpose of comparison, the tire model was adapted to fit the tire model in the multibody software. However, it should be noted that the tire configurations were not the same, and this might have caused some differences in the results obtained from the two different programs. Simulations were performed to examine the vehicle\u2019s behavior when traveling at a constant speed on a road with a single obstacle that had a rigid radius of 200 mm. Fig. 9 shows the side view of the vehicle when the front wheels passed over the obstacle. The results presented here describe the travel of the vehicle at a velocity of 5 m/s. In all the illustrations, DBD is compared with simulation of multibody dynamics using the Siemens\u2019 VL software. The kinematics relative to the horizontal position of the vehicle are described in Fig. 10. The illustration on the right describes the position of the center of gravity of the chassis on the Z axis, and the illustration on the left describes the pitch angle (negative value describe pitch up according to the reference system as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000939_s12206-020-0434-7-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000939_s12206-020-0434-7-Figure6-1.png",
        "caption": "Fig. 6. Contact and deformation before loading.",
        "texts": [
            " To facilitate comparison and analysis, the traditional high-speed ball bearing analysis model that accounts for the balls\u2019 load and the relative position of the centers of the ball and the centers of curvature of the inner and outer raceway grooves are illustrated in Fig. 4. According to the relative position relationship between ball center and curvature center of inner and outer grooves at any position, the distance between the ball center to curvature center of inner and outer grooves can be regarded as a virtual bar respectively are illustrated in Fig. 5. Before loading, the geometric properties of the ball bearing are depicted in Fig. 6. Because =e e wR f D , the distance between the center of curvature of the fixed outer raceway groove and the center of the ball can be given as: w w/ 2=( 0.5)= - -ej e el R D f D (1) Similarly, for the inner raceway groove: w( 0.5)= -ij il f D . (2) Further, the distance between the centers of curvature of the inner and outer raceway grooves can be estimated using the following equation: w( 1) .= + -i eA f f D (3) According to the relationship between the load along the virtual bar direction ejQ and the ijQ elongation of the virtual bar system after loading in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000939_s12206-020-0434-7-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000939_s12206-020-0434-7-Figure5-1.png",
        "caption": "Fig. 5. Rod model.",
        "texts": [
            " To facilitate comparison and analysis, the traditional high-speed ball bearing analysis model that accounts for the balls\u2019 load and the relative position of the centers of the ball and the centers of curvature of the inner and outer raceway grooves are illustrated in Fig. 4. According to the relative position relationship between ball center and curvature center of inner and outer grooves at any position, the distance between the ball center to curvature center of inner and outer grooves can be regarded as a virtual bar respectively are illustrated in Fig. 5. Before loading, the geometric properties of the ball bearing are depicted in Fig. 6. Because =e e wR f D , the distance between the center of curvature of the fixed outer raceway groove and the center of the ball can be given as: w w/ 2=( 0.5)= - -ej e el R D f D (1) Similarly, for the inner raceway groove: w( 0.5)= -ij il f D . (2) Further, the distance between the centers of curvature of the inner and outer raceway grooves can be estimated using the following equation: w( 1) .= + -i eA f f D (3) According to the relationship between the load along the virtual bar direction ejQ and the ijQ elongation of the virtual bar system after loading in Fig",
            " According to the principle of force transfer, the force acting on the rigid body can be moved to any point along its line of action without changing the effect of the force on the rigid body. Based on this principle, the normal load between the inner raceway groove and the ball can be translated to the center of curvature of the inner raceway grooves along the line connecting the inner center of curvature and the steel ball; further, the normal load between the outer raceway groove and the steel ball can be translated to the center of curvature of the outer raceway grooves along the line connecting the outer center of curvature and the steel ball, as depicted in Fig. 5. According to the line-of-force translation theorem, the force acting on a point of the rigid body can be moved to any parallel point; however, a coupled force must be added, and the moment of this additional coupled force is equal to the moment between the new force point and the original force. Using this theorem, the friction acting on the steel ball can be translated to the centers of curvature of the inner and outer raceway grooves, and the corresponding coupled force can be obtained. In other words, the friction force acting on the steel ball from the outer raceway groove can be translated to the center of curvature of the outer raceway groove while the additional coupled force 1jM is attached; further, the friction on the ball is translated to the center of curvature of the inner raceway groove with the additional coupled force 2jM , as depicted in Fig. 5. The aforementioned moments of force 1jM and 2jM can be given as follows: 1j 1j =( ( 0.5) )2l l= + - + D gjw ej gj w e w ej ej w MDM L M D f D l D (8) 2 j 2j =( ( 0.5) )2l l= + - + D gjw ij gj w e w ej ej w MDM L M D f D l D (9) Further, the combined moment acting on the rods model can be given as follows: 1 2 (( 0.5) ) (( 0.5) ) 0 = + - + D + - + D - - = gj gj ej e w ej w gj ij i w ij j j w M M M f D l D M f D l M M D l l (10) In the formula: gjM denotes the gyroscopic moment, 12 54.47 10 sin b-= \u00b4gj w Rj mj jM D n n ; lej and ,ijl respectively, represent the outer and inner ring channel control parameters of the bearing; Rjn indicates the thj ball revolution speed; b j represents the thj ball rotation attitude angle",
            "5) cos ( 1) cos \u00e9 \u00f9 \u00e9 \u00f9- + D + - + D\u00eb \u00fb \u00eb \u00fb = + - + i w j ij e w j ej i e w o r ij f D l f D l f f D a a a d (13) According to the axial displacement of the center of curvature of the inner raceway groove, the following geometrical relation holds: 2 1 _ ( 0.5) sin ( 0.5) sin ( 1) sin \u00e9 \u00f9 \u00e9 \u00f9- + D + - + D\u00eb \u00fb \u00eb \u00fb = + - + i w j ij e w j ej i e w o a ij f D l f D l f f D a a a d (14) In the formula: a ij , ,eja respectively, represent the actual contact angle of the ball with the inner and outer rings. 0a represent the initial contact angle. The simultaneous Eqs. (11)-(14) obtain the deformation coordination equation of the virtual rod system: As depicted in Fig. 5, the following equations can be obtained by considering the balance of horizontal and vertical forces. According to the contact load of the outer and inner ferrules of the ball and the load of the inner outer ring shared by the gyro moment, as shown in Fig. 5, the following equations can be obtained by considering the balance of horizontal and vertical forces: sin sin ( cos cos ) 0a a l a l a- - - =gj ij ij ej ej ij ij ej ej w M Q Q D (16) In the formula: cjF represent the centrifugal force acting on the ball, 3 2 12r w=cj w mF D D , r represent the material density of the ball, w is the angular velocity of the rolling element. Based on the balance equation of virtual bar system in horizontal and vertical directions, the force balance equation of virtual bar is determined: cos sin sin cot cos cot 0 + + - + = ij gj ij ij cj ij ij ij ej ij gj ij ej M Q F Q D M D l a a a a l a a (18) According to the coupling relationship between ijQ , ejQ and the whole bearing load, find the balance between the entire bearing and the external force: 1 ( sin cos ) 0 l a a = = - - =\u00e5 j z ij gj a ij ij ij j w M F Q D (19) 1 ( cos sin )cos 0 l a a y = = - - =\u00e5 j z ij gj r ij ij ij j j w M F Q D (20) 2 1 [( sin cos ) ]cos 0 2 l l a a y = = - - \u00c2 + + =\u00e5 j z ij gj ij m gj ij ij ij i j j j w w M D M M Q M D D (21) In the formula: aF represent the axial force of the bearing, rF represent the radial force of the bearing, j represent the rolling element at the angular position, Z represent the number of rolling elements, y j represent the angular position of the rolling element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure41.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure41.6-1.png",
        "caption": "Fig. 41.6 CAD Exploration steps of PIPV air blower design",
        "texts": [
            " Once the dimensions of all the components are fixed, the data is given to the industrial designer instead exploring the form on sketching designer preferred to model the provided components onCAD to get actual visualization and iteration of arrangement of various elements. Below are the steps recorded from the CAD. CAD gives freedom for designer to arrange the components in required manner. After basic arrangement of components, designer tried iterative sketches to improve on the form of product. It started with the modelling of panel followed by modelling of functional components like battery, fan, etc., then exploration of product body form and in last fine tuning of the form as shown in Fig. 41.6. Designer considered various issues of usability, aesthetics and functionality of product while exploring the form. Figure 41.7 shows the concept presentation by designer; product is named as JHONKA as it is the word for breeze in Hindi. Hundred PIPV product are identified from various platforms including online marketing, retail shops from Guwahati and Pune city. Information about Innovations and attempts to innovate is gathered from \u2018Techpedia\u2019 an online platform provided by National Innovation foundation and various individual designers working in this sector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002323_978-3-030-67411-3_32-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002323_978-3-030-67411-3_32-Figure17-1.png",
        "caption": "Fig. 17. Centrifugal force model and its operation.",
        "texts": [
            " When using a rope (string) that is tight to a flexible end, it is possible to tune the base frequency or higher harmonic frequencies, measure the speed of the wave travel in the rope and tune the frequencies for different number of nodes in a static wave. The pulses to the rope are generated by an oscillating mechanism (Figs. 15 and 16). A moving body would continue in a straight motion unless we apply a force. If it is bound by a string, or some other way it will move along a circular trajectory, but it will pull the string with a centrifugal force. How large is this force? Pupils measure the force using the force sensor in this exercise and visualize the results in chart again (Fig. 17). Spike Up Prime Interest in Physics 157 We have tested some of the experiments in the elementary school robot club. We have found that the children felt very comfortable using the building instructions, and had fun building the models. They were also confidently downloading and running the programs for the experiments. In the case of the scales experiment, we have discovered a small discrepancy in the scales model resulting in different measured values, but we have corrected the model based on this experience"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000048_j.biosystemseng.2019.10.007-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000048_j.biosystemseng.2019.10.007-Figure1-1.png",
        "caption": "Fig. 1 e Different rings of a MEA and a central disc.",
        "texts": [
            " However, a tensile force might deform the MEA in the axial direction, absorbing some rollover energy of the vehicle and contributing to the rotation of the ROPS around the opposite side, preventing the infringement of the safety volume for this cause. In this paper, a MEA is designed containing a sequence of solid rings and a central disc. Between these elements, some arrays of arms have been created. The arms rotate, as the MEA deforms in the axial direction. The displacement of the central disc with respect to the outer ring is performed thanks to the rotation of the arms. In Fig. 1 aMEAwith four rings, in addition to a central disc, is displayed. The number of arms in each array can vary, depending on the design. The smaller the number of arms in an array, the greater the potential displacement between the rings or ring and central disc could be reached. Figure 2 shows the different arrays of arms in a specific MEA and an example of an arm in each array. In the design of a MEA, it is possible to make many decisions in relation with the details of its geometry, as well as on its manufacturing process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002570_iros.2011.6048138-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002570_iros.2011.6048138-Figure13-1.png",
        "caption": "Fig. 13. Experimental setup of press force adjustment mechanism",
        "texts": [
            " The motor does not rotated when the clearance is from 0 to 0.1 mm. The vibration source of this motor is a 26.8 kHz Langevin transducer placed on the end of the waveguide. This transducer has no control function for its power, but is able to control the efficiency of vibration propagation by adjustment of the press force from the vibration horn to the waveguide. To investigate the relationship between press force and rotation speed or starting torque, the rotation test is executed. As shown in Fig. 13, the waveguide is set between the ultrasonic vibrator horn and the buffer material made of silicone. The press force is adjusted by movement of the Z-axis stage below the buffer material. The operator adjusts the zero position between the vibration horn and the waveguide by visual observation, it is set when the horn touches the waveguide upon manual adjustment of the Zaxis stage. The inner diameter of the rotor is 0.91 mm. Press force is measured by the digital force gage set between the stage and the buffer material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000249_icems.2019.8921452-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000249_icems.2019.8921452-Figure4-1.png",
        "caption": "Fig. 4. Flux line distributions. (a) Excited only by stator PMs. (b) Excited only by rotor PMs. (c) Excited only by armature windings.",
        "texts": [
            " It can be found that there are three groups of dominant harmonics in Table I, as long as it satisfies 3 1 2p p p= \u2212 (4) ( )3 1 3 1p p\u03a9 = \u03a9 (5) These effective harmonics in each group have the same PPN and rotational speed. So the stable electromagnetic torque can be achieved by the interaction among these effective harmonics. B. Verification of Working Principle Magnetic field analysis by FEM is conducted to verify the working principle. In this part, a HS-DPME motor with p1=10, p2=9 and Zs =12 is selected for analysis. Fig. 4 illustrates the flux line distributions excited only by stator PMs, rotor PMs and armature windings, respectively. As can be seen, although the PPNs of the three field sources are different from each other, the PPNs of the modulated magnetic fields are identical as reflected in the stator yoke and rotor yoke. Fig. 5 exhibits flux density waveforms in airgap produced by the three field sources and corresponding harmonic spectrum. As can be seen from the harmonic spectrum, PPNs in rectangles agree well with the foregoing theoretical analysis in Table I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure15-1.png",
        "caption": "Fig. 15 Configuration 2",
        "texts": [
            " When the planar double-folded metamorphic mechanism is in configurations 1 and 3, the mechanism under force constraint can be regarded as consisting of an active part and an augmented Assur group RRRP (as shown in Fig.\u00a012b and c). The dynamic analysis of the planar double-folded metamorphic mechanism in configurations 1 and 3 is shown in Fig.\u00a014. The dynamic equations in configurations 1 and 3 can be obtained by Eqs.\u00a0(1) and (4) as follows, When the mechanism is in configuration 2, the mechanism under geometric constraint can be regarded as consisting of an active part and an Assur group RRR, as shown in Fig.\u00a015. The dynamic analysis of the planar double-folded metamorphic mechanism in configuration 2 is shown in Fig.\u00a014a and b. The dynamic equations in configuration 2 can be obtained by Eqs.\u00a0(1) and (6) as follows, (11) \u23a7\u23aa\u23a8\u23aa\u23a9 Fax + Fbx = m1a1x Fay + Fby \u2212 G1 = m1a1y 0.5LABFbx sin(\ud835\udf031 \u2212 \u03c0) + 0.5LABFby cos(\ud835\udf031 \u2212 \u03c0) +M1 \u2212 0.5LABFax sin(\ud835\udf031 \u2212 \u03c0) \u2212 0.5LABFay cos(\ud835\udf031 \u2212 \u03c0) = J1?\u0308?1 (12) \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23a9 \u2212Fbx + Fdx = m2a2x + m3a3x \u2212Fby + Fdy \u2212 G2 \u2212 G3 = m2a2y + m3a3y \u2212LBCFbx sin \ud835\udf032 + LBCFby cos \ud835\udf032 + 0.5LBCG2 cos \ud835\udf032 = JBC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000305_iecon.2019.8927454-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000305_iecon.2019.8927454-Figure4-1.png",
        "caption": "Fig. 4. Test bench for the dual three-phase PMSM.",
        "texts": [
            " Note that in (20) the rotor angle \u03d5 is mapped to the interval [\u2212\u03c0, \u03c0] by means of a modulo operation. The identification task can then be formulated as finding optimal values of the Fourier coefficients \u03b1n, \u03b2n, such that the (squared) error between a measured torque \u03c4m and the simulated torque \u03c4 is minimized. For this purpose, measurements of the currents and torque for a number N\u03c4 i = 6 constant terminal currents1 and for N\u03c4 \u03d5 = 180 fixed rotor angles (step size \u2206\u03d5 = 0.5\u00b0) within one electric period of the PMSM of 90\u00b0 were performed on the test bench shown in Fig. 4. Starting from the left, the considered PMSM is coupled with a rotary encoder, followed by a torque sensor, a second rotary encoder, fly wheels and a load machine. The resulting (nonlinear) optimization problem can be formulated as min x N\u03c4\u03d5\u2211 j=1 N\u03c4i\u2211 k=1 q ( \u03c4jk \u2212 \u03c4mjk )2 (22a) subject to 0 = gjk, j = 1, . . . , N\u03c4 \u03d5, k = 1, . . . , N\u03c4 i (22b) with q > 0, see also [12]. The equality constraints (22b) result from the magneto-static model (17b) evaluated at each measured current iI,mcL,jk and angle \u03d5mjk gjk = ( Gt,jk + DgGc,jkD T g ) utg,jk + DgGc,jk ( D\u0303T c H I mLi I,m cL,jk + DT mutm ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001707_0309324720958257-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001707_0309324720958257-Figure2-1.png",
        "caption": "Figure 2. Schematic diagram of tension setting out.",
        "texts": [
            "raided wire rope, Mechanical model, Mechanical response, Test verification Date received: 23 March 2020; accepted: 14 August 2020 A braided wire rope is made of a set of left-handed turns and a set of right-handed single-stranded strands of wire ropes that are woven by intersecting spiral tracks. The left-handed and right-handed strands are equal in number and woven symmetrically. The geometric model is shown in the Figure 1. The anti-torsion characteristic of a braided wire rope is widely used as a guide rope and a traction rope in the tension line of the transmission line, as shown in Figure 2. The construction mode determines that the braided wire rope mainly bears the tensile load. However, due to the uneven stress of the strands, different deformations occur and the original balanced structure gets destroyed, resulting in the internal torsion of the transverse rotation caused by the tension. Because the cross section of a braided wire rope is non-circular (quadruple or hexagonal), the torsional load from the outside is generated when the wheel grooves such as the tractor or the pulley have frictional contact during the wire laying process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002993_s12206-021-0629-6-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002993_s12206-021-0629-6-Figure1-1.png",
        "caption": "Fig. 1. Generation process and mathematical modeling system of worm.",
        "texts": [
            " Then, the matched worm wheel is generated by the formed worm surface. This special worm and worm wheel surface are typically complex surfaces in engineering. The meshing optimization research in this study focuses on the whole meshing cycle and all meshing lines, as a result, the tooth surface of the worm or worm wheel should be calculated and obtained at first. In this study, worm surface is taken as an instance for meshing optimization, thus mathematical model construction of the worm is required. The mathematical modeling coordinate system is built as Fig. 1, where a is the center distance of worm drive; \u03b2 is the tilt angle of the mother plane; \u03c62 and \u03c63 are angular displacements for the worm and tool rest (also worm wheel), respectively; and db is the circle diameter of the tool rest (also base circle diameter of worm wheel). Related coordinate systems are defined and illustrated as follows: \u00b7S1(O1-xyz) refers to the mother plane coordinate system. \u00b7S2(O2-xyz) refers to the static worm coordinate system, and also represents the initial envelope movement of worm",
            " \u00b7S3(O3-xyz) refers to the static tool rest coordinate system, also represents the initial envelope movement of tool rest. \u00b7S3i(O3-xiyizi) represents the moving coordinate system of tool rest. Coordinate of a meshing point in S1(O1-xyz) is described as [x1, y1, z1]T, coordinate in S2(O2-xyz) is indicated as [x2, y2, z2]T, coordinate in S2i(O2-xiyizi) is named as [x2i, y2i, z2i]T, coordinate in S3(O3-xyz) is represented as [x3, y3, z3]T, coordinate in S3i(O3xiyizi) is described as [x3i, y3i, z3i]T. According to the mathematical modeling system demonstrated in Fig. 1, although [x1, y1, z1]T is given, the meshing point in S3(O3-xyz) could be calculated by using coordinate transformation matrix: 3 1 3 1 13 0 ( ) * ( ) * 2 2 2 0 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5= \u2212 \u2212 + \u2212\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6\u23a3 \u23a6 \u23a3 \u23a6 b y z x x dy Rot Rot y zz \u03c0 \u03c0\u03b2 (1) where Rotz(-\u03c0/2) is the coordinate transformation matrix that rotates around z-axis, and -\u03c0/2 is the rotation angle. Other coordinate transformation matrices are homologous with this. Given that the mother surface is a plane, z1 = 0, x1 and y1 are variables that need to be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002987_s40964-021-00199-x-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002987_s40964-021-00199-x-Figure16-1.png",
        "caption": "Fig. 16 Orientation resulting from GA optimisation",
        "texts": [
            " This finding suggests that using a lower Ns would have only marginally affected the final solution. As an example, if stopping the calculation after 5 stagnating iterations, the algorithm would have been terminated in 9 iterations (instead of 19) with a worsening 1 3 of the fitness function equal to 0.02% . However, this result cannot be generalised as it depends on the specific combination of geometry and part objectives. The best chromosome at the end of the algorithm corresponds to Euler angles x = 262\u25e6 , y = 349\u25e6 , and z = 81\u25e6 . Figure\u00a016 illustrates this orientation. The part oriented as proposed by the system is manufactured using the same process parameters as detailed above. Figure\u00a017a shows the map of stair stepping defect in the proposed orientation. It is possible to notice that the calculated value of dst is less than 0.25 for all the elements of the cluster 1 in Fig.\u00a014 a. Specifically, dst is equal to 0.043 on the roof, 0.170 on the front deck, and between 0.001 and 0.233 on the rails of the boat. The comparison of these results with the ones presented in Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure2-1.png",
        "caption": "Fig. 2. Schematic representation of UPR and PRU limbs: (a) UPR limb; (b) PRU limb.",
        "texts": [
            " In the following diagrams, the fixed base is hidden to achieve a clearer expression. 2.2 Mobility analysis Screw theory [31] is used to analyze the mobility of the 2UPR-PRU PKM and the inverse dynamics in the following sections. The coordinates of points iA (i = 1, 2, 3) relative to O-xyz can be expressed as ( )T i i iA A Ax y z . The coordinates of iB (i = 1, 2, 3) relative to O-xyz can be expressed as ( )T 20 0l , ( )T 20 0\u2212l , and ( )T 3 0 0\u2212q , respectively, where 3q denotes the distance between points 3B and O. Fig. 2(a) shows that in the O-xyz, the twist system of the first UPR limb can be expressed as: ( ) ( ) ( ) ( ) 1 1 1 1 1 1 1 11 12 2 2 13 2 14 0 1 0; 0 0 0 c 0 s ; s 0 c 0 0 0; c 0 s ; s c s c \u23a7 = \u23aa \u23aa = \u2212\u23aa \u23a8 = \u2212\u23aa \u23aa = \u2212 \u2212 + \u2212\u23aa\u23a9 A A A A A A A l l x y l z y z x y \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ $ $ (1) where ij$ denotes the twist of the jth joint in the ith limb and \u03b2 is the rotational angle around the y-axis. Given the structural constraints of two UPR limbs, the coordinate iA x is equal to tan iA z \u03b2 (i = 1, 2)",
            " Adopting the reciprocal screw theory, the limb wrench system can be expressed as: ( ) ( ) 11 2 2 12 c 0 s ; s 0 c 0 0 0; s 0 c \u23a7 = \u2212 \u2212 \u2212\u23aa \u23a8 =\u23aa\u23a9 r r l l\u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ (2) where r ij$ denotes the jth unit constraint wrench of limb i. In Eq. (2), the constraint force 11 r$ passes through 1B , and is along the axis of R joint, and constraint couple 12 r$ is perpendicular to the U joint axes. Clearly, these two wrenches are reciprocal to all the twist screws of the first UPR limb. The wrench system of the second UPR limb is similar to that of limb 1. Fig. 2(b) shows that based on the distributions of twist screws of limb 3 (PRU), the wrench systems of limb 3 can be obtained easily as: ( ) ( ) 3 331 32 0 1 0; tan 0 0 0 0 0; s 0 c \u23a7 = \u2212 \u2212\u23aa \u23a8 =\u23aa\u23a9 r A A r x z\u03b2 \u03b2 \u03b2 $ $ (3) where tan denotes the tangent function. Based on Eqs. (2) and (3), the overall constraint wrenches can be expressed as: ( ) ( ) ( ) 3 3 1 2 3 c 0 s ; 0 0 0 0 1 0; tan 0 0 0 0 0; s 0 c \u23a7 = \u2212 \u23aa \u23aa = \u2212 \u2212\u23a8 \u23aa =\u23aa\u23a9 C C A A C x z \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ $ . (4) The overall motion of the proposed mechanism can thus be expressed as: ( ) ( ) ( )3 3 1 2 3 0 0 0; s 0 c 0 1 0; 0 0 0 c 0 s ; 0 c s 0 \u23a7 = \u23aa\u23aa =\u23a8 \u23aa = \u2212 +\u23aa\u23a9 pm pm pm A Az x \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ $ ",
            " (12) The mapping relationship of two velocities, pt and \u03b7 can thus be written as: TT T\u23a1 \u23a4= =\u23a3 \u23a6p p p Mt \u03c9 v J \u03b7 (13) where pv and p\u03c9 denote the translational and rotational velocities of moving platform, respectively, and the MJ is: T2 T , , 0 1 0 sec 0 0 c 0 s 0 0 0 0 0 0 tan 0 1 \u23a1 \u23a4 \u23a2 \u23a5\u23a1 \u23a4= = \u2212\u23a2 \u23a5\u23a3 \u23a6 \u23a2 \u23a5 \u23a3 \u23a6 o M R M T M z \u03b2 \u03b2 \u03b2 \u03b2 J J J (14) where the matrices ,R MJ and ,T MJ denote the rotational and translational parts of MJ , respectively. Through differentiation of Eq. (13), the accelerations of the moving platform can be obtained as: = +p M Mt J \u03b7 J \u03b7 (15) where T \u23a1 \u23a4= \u23a3 \u23a6oz\u03b2 \u03b3\u03b7 , and T 2 2 2 0 0 0 sec 2 tan sec 0 0 s 0 c 0 0 0 0 0 0 sec 0 0 \u23a1 \u23a4+ \u23a2 \u23a5 = \u2212 \u2212\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 o o M z z\u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 J . 4.2 Velocity and acceleration of limbs Here, the first UPR limb is selected as an example to introduce the derivation of velocity and acceleration analysis of limbs. Fig. 2(a) shows that through linear combination of twist screws [33] in limb 1, the velocity of moving platform can be obtained and expressed as: ( ) 1 TTTT T T 11 11 12 12 13 13 14 14 T 11 12 13 14 \u23a1 \u23a4\u23a1 \u23a4= = + \u00d7\u23a3 \u23a6 \u23a2 \u23a5\u23a3 \u23a6 = + + + = \u23a1 \u23a4\u23a3 \u23a6 o o o p p p O o k \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 V \u03c9 v \u03c9 v \u03c9 r $ $ $ $ J (16) where ov denotes the translational velocity of a point in the moving platform that is coincident with the origin point O, O or denotes the position vector from point O to o, 1 j\u03c9 (j = 1, 2, 3, and 4) denotes intensity of joint velocity in limb 1, and 1k J denotes a 6 4\u00d7 kinematic Jacobian matrix",
            " (24) Similarly, the acceleration of the jth joint in the limb 1, 1jA , can be written as: ( ) TTTT T T 1 1 1 1 1 1 1 11 11 12 12 1 1 ,1\u0394 \u23a1 \u23a4\u23a1 \u23a4= = \u2212 \u00d7\u23a3 \u23a6 \u23a2 \u23a5\u23a3 \u23a6 = + + + + j j j j j j j j j j\u03c9 \u03c9 \u03c9 A a v v $ $ $ $ \u03c9 \u03c9 \u03c9 (25) where ( ) ( ) ,1 11 11 12 12 1 1 12 12 13 13 1 1 1 11 1 1 1+ + \u0394 \u2212 \u2212 \u23a1 \u23a4= + +\u23a3 \u23a6 \u23a1 \u23a4\u23a1 \u23a4+ +\u23a3 \u23a6 \u23a3 \u23a6 j j j j j j jj j \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 $ $ $ $ $ $ $ $ $ . The acceleration of the center of mass of the jth link in limb 1, 1 mj,C A , can then be written as: ( )( ) TT T 1 , 1 1 , TT T 1 1 1 1 , 1 1 1 , \u23a1 \u23a4= \u23a3 \u23a6 \u23a1 \u23a4= + \u00d7 + \u00d7 \u00d7\u23a2 \u23a5\u23a3 \u23a6 m m m m j C j j C j j j j C j j j C A a v r r \u03c9 \u03c9 \u03c9 \u03c9 \u03c9 . (26) Fig. 2(b) shows that the velocity and acceleration results of PRU limb also can be obtained using the above derivation. 4.3 Actuated force analysis In this study, the actuation forces of 2UPR-PRU PKM are calculated using the principle of virtual work. The force/torque acting on the center of mass of the moving platform can be generally written as: T , , , , , ,( )\u23a1 \u23a4= \u2212 + \u2212 \u2212 \u00d7\u23a3 \u23a6m m m m O O p C p p C p E C p E C p p p p pmF g a f I I\u03c4 \u03c9 \u03c9 \u03c9 (27) where pm denotes the moving platform\u2019s mass; g denotes gravity; , mp Ca denotes the linear acceleration of the center of mass of the moving platform; O pI denotes the inertia matrix of the moving platform relative to O-xyz, which can be obtained by combining the rotation matrix O oR and the inertia matrix of the moving platform relative to the o-uvw; and , , mp E Cf and , , mp E C\u03c4 denote the external force and torque acting on the center of mass of the moving platform, which are set as zero in this study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001859_j.promfg.2020.10.106-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001859_j.promfg.2020.10.106-Figure1-1.png",
        "caption": "Fig. 1. Powder delivery piston with recoater (doctor) blade in SLS machine.",
        "texts": [
            " Hence, little space is allowed for researchers who desire to experiment with powder dynamics and rheology parameters to examine their effect on the quality of the deposited nonsintered powder layer. Limitations in controlling both powder spreading and laser beam parameters is the main reason for development of open-system SLS machines, to examine parameter correlations that could not have been examined in the closed, industrial SLS machines [9]. In order to achieve homogeneous, high quality powder layers, various custom powder deposition systems have been designed. When it comes to raw powder storage, the most common solution is the powder delivery piston, see Fig. 1. However, this means that the powder that is stored must be properly sieved beforehand, so that is guaranteed to be clear of agglomerates and have the desired grain size distribution. Since humidity of the environment can cause agglomerate formation, the powder storing tank above the piston must be t . li l i t . i i ti l t li tt :// ti . /li / / . / i i ilit t i ti i itt t . , . . , , , , , , , . , . , . . \u00a9 . . . . w . l ti i t i ; iti t ; ll ; i i ; t 756 Panagiotis Avrampos et al. / Procedia Manufacturing 51 (2020) 755\u2013762 2 Author name /Procedia Manufacturing 00 (2019) 000\u2013000 emptied and refilled regularly, otherwise the conditions of the raw powder can be considered neither stable nor known",
            " The powder hopper/doser subsystem consists of the powder tank in which powder is stored and the doser drum, which is a cylinder with a helical pattern of blind holes of fixed volume that is rotated by a stepper motor, see Fig. 2. The powder tank can be sealed and equipped with humidity and temperature sensors in order to measure the environmental factors that actively affect the powder quality. Panagiotis Avrampos et al. / Procedia Manufacturing 51 (2020) 755\u2013762 757 2 Author name /Procedia Manufacturing 00 (2019) 000\u2013000 emptied and refilled regularly, otherwise the conditions of the raw powder can be considered neither stable nor known. Fig. 1. Powder delivery piston with recoater (doctor) blade in SLS machine. Commercial PDS come with a roller and/or a doctor blade, as means of applying an evenly spread powder layer. The doctor blade, however, only scrapes the powder, hence it is referred to as scraper as well, thereby hardly providing any powder compaction [10]. By contrast, a counter-rotating roller does compress the powder and spread an even layer. The compaction level depends on the height of the non-compacted powder that comes to contact with the cylinder compared to the height of the compacted powder that the cylinder leaves behind"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001787_ecce44975.2020.9235832-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001787_ecce44975.2020.9235832-Figure1-1.png",
        "caption": "Fig. 1. (a) Six-pole hybrid rotor-excited synchronous machine [5] and (b) eight-pole synchronous machine with proposed parallel hybrid-excited rotor.",
        "texts": [
            " Downloaded on June 24,2021 at 15:17:54 UTC from IEEE Xplore. Restrictions apply. The most cited paper on the topic of hybrid excitation is the paper of [5]. Yet it is interesting that no further attention has been paid in the published research to this parallel hybrid excitation method. The method is also not investigated in any detail regarding the effect of magnetic saturation. In this section, we investigate this proposed method further. The cross-section of the hybrid-excited rotor synchronous machine proposed by [5] is shown in Fig. 1(a). The generator has 36 slots and six poles, of which four poles are PM poles and two poles are field-winding poles. It is clear in this case that the maximum number of parallel circuits of the stator is two, since a series-connected phase group must span across two PM poles and one field-winding pole. The latter is important because, if the generator was just an ordinary six-pole PM generator, then up to six parallel circuits could be used. In this regard it already points out the disadvantage of this method of parallel hybrid excitation, an aspect that is further considered later in the paper. When the effect of magnetic saturation in the magnetic equivalent circuit of the machine in Fig. 1(a) is taken into account, it is clear that the large stator yoke reluctances that magnetically link the two field-winding poles disturb the parallel, independent action between the magnet and field excitations. This means that a much larger change in the field MMF is needed to get a certain change in flux linkage and induced voltage, which impairs the efficiency of the generator. Finally, a comment must be made about the use of negative field current in the machine in Fig. 1(a). This is proposed by [5] to get a wider variation in flux. With negative field current, the field winding poles swap which in effect changes the rotor to a two-pole rotor, as indicated by the north-south indications on the rotor in Fig. 1(a). A two-pole rotor per se is not a problem, because there is still a six-pole stator winding. However, with two-pole flux, the small six-pole designed stator yokes will undergo severe saturation unless they are designed to be large, which increases the mass and is uneconomical. Hence, in this paper the use of negative field current is not followed. In order to minimise the problem of the effect of magnetic saturation in the machine of Fig. 1(a) and to obtain an improved parallel hybrid excitation, the following requirements are set with respect to the layout of the hybrid rotor of integral, overlap stator-winding synchronous machines: (i) A minimum of two field-winding poles (pf ) must always be placed adjacently on the rotor. This is to shorten the stator yoke part that magnetically links the two field winding poles. To avoid unbalanced magnetic pull in the machine, a further minimum of two adjacent fieldwinding poles must be located on the opposite side of the rotor, thus displaced 180\u25e6 apart mechanically",
            " (ii) Similarly, a minimum of two PM poles (pm) must always be placed adjacently on the rotor and therefore there also should be at least two additional, adjacent PM poles on the opposite side of the rotor. This requires that there should also be a total of at least four PM poles. With these requirements, it means that the generator must have atleast eight poles or more. Therefore, for an eight-pole generator (p = 8), only one layout will be possible, namely four field poles (pf = 4) and four PM poles (pm = 4), as shown in Fig. 1(b). One can see in Fig. 1(b) how two field-winding poles are together and how two magnet poles are together. For a 12-pole generator, there will be two layout options, namely six field poles and six PM poles, or four field poles and eight PM poles. In general, for overlap stator-winding machines, we thus have the hybrid pole conditions as given by (1). Note in (1) that pm pf because the machine is mainly a PM machine, with only some additional flux variability by means of the field winding. The maximum number of parallel circuits, amax, can be determined for integral overlap stator windings by (2). p = pf + pm 8 pf 4; pf = even pm 4; pm = even pm pf (1) amax = 2nmax with nmax = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 n = 1, 2, 3, ... pf 2n = 2, 3, 5, 7, ... n \u2192 maximised \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 (2) For non-overlap stator-windings, the above-mentioned saturation problem of the machine in Fig. 1(a) is, for all practical reasons, not an issue. The reason for this is that almost all nonoverlap windings have two or more north-south adjacent poles in a series-connected winding section. Thus, for non-overlap winding machines, we need to consider only winding-pole sections (Ws) instead of number of poles. Hence, we divide the machine into a number of wound-pole sections (Wf ) and magnet-pole sections (Wm). For balanced magnetic pull, the rotor must have at least two wound-pole sections on opposite sides of the rotor and two magnet-pole sections on opposite sides of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000149_1350650119887043-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000149_1350650119887043-Figure1-1.png",
        "caption": "Figure 1. Schematic of the radial bump foil bearing.",
        "texts": [
            "1 Theoretically, foil bearings do not have any speed limitations. Foil bearings differ based on the type of elastic foundation used. Elastic support can be in the form of cantilever, spiral, bump, or wire mesh. In this work bump type foil bearings are considered as they are most commonly used foil bearing due to their simpler design and ease of fabrication to suit any size. The radial bump foil bearing consists of smooth encircling or top foil supported on a layer of bump foils contained in a rigid bearing housing, as shown in Figure 1. Both the encircling foil and bump foil are fixed at one end and another end is set free. This permits the deformation of flexible support under aerodynamic pressure and facilitates air film formation. In stationary condition, there exists a small clearance of the order of few microns between the shaft and smooth top foil. As the shaft starts rotating, pressure develops between the shaft and the top foil due to aerodynamic action. As the pressure force exceeds the weight of the shaft, it gets completely airborne"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000017_chicc.2019.8866432-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000017_chicc.2019.8866432-Figure10-1.png",
        "caption": "Fig. 10 Water quality sensor delivery platform. As shown in Fig. 10, the STM32 board control the motor's forward and reverse and the motor drives the conveyor belt and the fixed pulley to realize the function of sensors\u2019 delivery and recovery. The rotation speed and running time of motor is fixed to ensure the same drop height of the sensor during each task.",
        "texts": [
            " If the flying distance of the drone is within 3 kilometers, we can use Lora module to acquire the water quality data or make a new plan for the drone. RS485 serial port is designed to be detachable to connect different sensors, so many types of data of water quality can be measured. One of the most challenging problem we encountered in the system was how to put the sensor into the water. In order to conquer the difficulty, a special platform is designed to lay down the sensors into the water. Fig. 10 illustrates the delivery platform. This paper presents a PID error compensated dynamic inversion controller for hexacopter carried with the IoT water quality measurement system. Newton-Eulerian\u2019s law is used for deriving the dynamic model of the hexacopter. Then we designed the position and attitude controller using cascade PID and DIC-PID respectively. The attitude controller is the focus of this paper which is divided into an inner and outer loop and the simulation are compared between the proposed DIC-PID controller and the traditional method without compensation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure6-1.png",
        "caption": "Fig. 6. Determination of the calculated COG.",
        "texts": [
            " Measurement 174 (2021) 108956 When the measured body is suspended on the pendulum, since the measured body and the pendulum are connected by a universal joint, in theory the COG of the measured body automatically falls on the pendulum axis. Under different suspended postures, the intersect point of pendulum axis is assumed as COG [5]. In measurement, the OdZd-axis of PCS is regarded as the pendulum axis, but due to measurement errors, the pendulum axes under different suspended postures do not absolutely intersect, and the calculated COG are obtained by the following method. As shown in Fig. 6, the lines l1 and l2 are the measured pendulum axes in RCS under two different suspended postures of the measured body respectively. Points Od1 and Od2 are the positions of the measured origins of PCS in the RCS under two suspended postures respectively. Assuming that vectors e1 and e2 are the direction vectors of lines l1 and l2 respectively, P1 and P2 are two points on lines l1 and l2 respectively, and line P1P2 is perpendicular to l1 and l2 simultaneously. The distances from the midpoint of the line P1P2 to the lines l1 and l2 is the smallest distance between the two lines, so the midpoint of the line P1P2 is regarded as the calculated COG which is represented by Ci"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003104_13506501211033045-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003104_13506501211033045-Figure1-1.png",
        "caption": "Figure 1. Fluid pivot bearing model.",
        "texts": [
            " Numerical model and dynamic characteristics of the pad To determine the dynamic characteristics of the pad, it is necessary to first determine the equilibrium position of the bearing-shaft system, including the balance of the shaft and that of the pads. The balance of the shaft in the bearingshaft systemhas to satisfy the hydrodynamic problem in the journal bearing. The balance of the pads has to satisfy the condition of flow balance and torque balance. Therefore, FPJB is not only more complicated than TPJB, but also its computation volume and calculation time are much larger. The geometrical model of FPJB is shown in Figure 1. Considering the static performance of bearing,Reynolds equation governing the steady, laminar incompressible flows in hydrodynamic oil film in dimensionless form (flow between the shaft and the inner surface of the pad): \u2202 \u2202\u03d5 H3 d \u2202Pd \u2202\u03d5 ( ) + D L ( )2 \u2202 \u2202\u03bb H3 d \u2202Pd \u2202\u03bb ( ) = \u2202Hd \u2202\u03d5 (1) And inhydrostatic oilfilm in thedimensionless form (the oil flow between the outside of the pad and the inside of the shell): \u2202 \u2202\u03d5 H3 s \u2202Ps \u2202\u03d5 ( ) + D L ( )2 \u2202 \u2202\u03bb H3 s \u2202Ps \u2202\u03bb ( ) = 0 (2) If it is assumed that there arenoplastic and thermal deformations of the shaft, pads, and shell, the hydrodynamic and hydrostatic oilfilm thickness is determined by the following formula: Hd,i = 1\u2212 mpcos(\u03d5MC,i \u2212 \u03d5)+ \u03b5cos(\u03d5\u2212 \u03b8) + Rpo\u03b1i Cp sin (\u03d5TC,i \u2212 \u03d5) (3) Hs,i = \u2212Rpo\u03b1i Cp sin (\u03d5TC,i \u2212 \u03d5) (4) At the early stage, under the influence of hydrodynamic pressure and hydrostatic pressure, the pad tends to move"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure6.11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure6.11-1.png",
        "caption": "Fig. 6.11 Schematic representation of a cold spray array",
        "texts": [
            " In such a process, small particles (~150\u00a0\u03bcm) are accelerated at a supersonic speed through the effect of a pressurized gas, which can be nitrogen or helium, and whose operating temperature reaches 800\u00a0\u00b0C.\u00a0The particles projected at supersonic 99 speed create a coating whose porosity tends to decrease with the successive layers of applied material, allowing to reach a high hardness and density. The use of inert gases and temperatures below the melting temperature in the applied materials decreases the degree of reactivity and therefore the presence of oxides in the coating while maintaining the initial properties of the projected material, as well as its microstructure (Fig.\u00a06.11). The quality of the coating varies depending on the conditions in which the cold projection is carried out. The parameters to consider are type of material projected, as well as its physical and chemical properties, type of substrate, the contact surface roughness, and operating conditions of the equipment used. The issue concerning the deposit parameters is especially important in the case of light alloys. In the case of aluminum alloys, the temperature and pressure conditions reported in the literature are 300\u00a0\u00b0C and 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002348_s12206-021-0133-z-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002348_s12206-021-0133-z-Figure1-1.png",
        "caption": "Fig. 1. The 2-(3-RRPS) parallel manipulator.",
        "texts": [
            " The experiments are performed for a 6-DoF 2-(3-RRPS) parallel manipulator proposed by Gallardo in Ref. [23]. The rest of this work is organized as follows: Sec. 2 presents a brief description of the manipulator under study. Sec. 3 analyzes the calibration process, and the Zhang\u2019s method. Sec. 4 introduces the measurement system proposed. Secs. 5 and 6 describe the methodology and its application for calculating the geometrical parameters of a robot prototype, respectively. Sec. 7 presents a case of study and, finally, Sec. 8 shows the conclusions of the paper. The robot under study, Fig. 1, was introduced by Gallardo [23], and it has the potential to serve as a positioning device for a multi-axis machine-tool. It consists of two 3-RRPS parallel robots sharing a common moving platform. Here R, P, and S stand for revolute, prismatic, and spherical joints, respectively. In this manipulator, the prismatic pair are the active joints. The mobile platform, Fig. 2, is a triangular prism whose upper and lower faces are equilateral triangles 1 2 3S S S and 4 5 6S S S , where the length of the sides is e , the height is h and point iS locates the center of the i -th spherical joint",
            " The method proceeds by performing several iterations according to the following mathematical scheme: ( ) 1 1 ( ) ( ) ( ) ( )T T s s s s s s \u2212 + = \u2212u u J u J u J u d u (8) where ( )J u is the Jacobian matrix defined by the partial derivatives of the function ( )id u with respect to each objective variable, ( ), , , Ta b c r , and vector ( )d u contains the functions ( )id u for all the data. It should be noted that parameter ir is also computed by the procedure. However, this geometrical parameter is not relevant for calibration purposes. On the other hand, point iS generates a spherical surface F iV , with center in iA . These points are located at the inter- section of the two axes of revolution of its corresponding chain, and further define the dimensions of the fixed platform, as shown in Fig. 1. Therefore, in order to determine the coordinates of points iA relative to the reference frame of the camera, an experiment is mounted as shown in Fig. 10. It is noteworthy that, for this procedure, it is not necessary to activate the actuator. Points C iA are determined in the same way that the coordinates of c iS . As shown in Fig. 2, the mobile platform is geometrically defined as the polyhedron whose vertices are the centers of rotation of the spherical joints. Following the methodology previously described, it is possible to locate these points since they correspond to the centers of the vertex spaces shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001226_acpee48638.2020.9136520-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001226_acpee48638.2020.9136520-Figure3-1.png",
        "caption": "Fig. 3. The diagram of effect analysis for conductivity defects length on the electric field distribution in single composite insulator string",
        "texts": [
            " 2, the position of the distance 0m is at the high voltage end of the insulator. The electric field distribution of the composite insulator string without the defect is U-shaped. For the 500kV transmission line composite insulator string, the electric field at the high voltage end of the insulator is 601.8kV/m kV/m, and it is 180.5kV/m at the ground end of the insulator. B. Influence of the Defect Length a) Defect in the Single Insulator String: The diagram of influence analysis for conductive defects length on the electric field distribution is shown in Fig 3. In Fig. 3, L is the length of the defect. The electric field distribution of composite insulator strings with conductive defect are shown in Fig. 4, and the length of the defect section was 0.6m. When the conductive defects exist in the insulator, the electric field distribution is no longer continuous. The insulator electric field distributions with the conductive defects of different lengths at the high voltage end are shown in Table II. In Table II, Eh is the electric field at the high voltage end of the insulator, Ed is the maximum electric field at the defect, and Eg is the electric field at the grounding end of the insulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001429_j.procir.2020.05.180-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001429_j.procir.2020.05.180-Figure2-1.png",
        "caption": "Fig. 2. Drawing of sample part with six planes and two cylinders. Recreated from [20].",
        "texts": [
            " This table may be the direct output from automatic feature recognition providing an overview of the constituent primitives of the part, as well as a starting point for tolerance analysis. Depending on the sophistication of the feature recognition module, the resulting table may include an unknown number of features not relevant for tolerancing purposes. Hence, the judgment of an engineer may be required to transform the automatically generated table to a suitable configuration. To demonstrate the approach described in the previous section, a simple part geometry recreated from [38] is used as a case study (Fig. 2). The STL file contains a list of 860 facets which may be reduced to the 8 constituent features as displayed in Table 4. The table directly enables the specification of the vectorial tolerances associated with the part features. To conduct a proper case study, the part should be assigned a function. Since the true purpose of the part is unknown, two assumptions are made about the design intent: \u2022 The component is a standardized part of an assembly where it is intended to connect two or more shafts; and \u2022 One possible application of the component requires support on the angled planes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000302_icaaid.2019.8934956-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000302_icaaid.2019.8934956-Figure1-1.png",
        "caption": "Fig. 1. TRMS description",
        "texts": [
            " Moreover, a new method for identifying the suggested nonlinear PID controller parameters has been introduced based on Differential Evolution (DE) algorithm which has the advantages of having few parameters, faster convergence and difficult to fall into the local optimum [23]. The rest of the paper is organized as follows. The next section gives the nonlinear description of TRMS. Section 3 introduces the mathematical relations of proposed controller. Following section explains how the linear and the novel nonlinear PID\u2019s parameters have been tuned by DE algorithm. Simulation and experimental results are provided in Section 4 and 5 respectively, followed by conclusions in Section 6. II. NONLINEAR MATHEMATICAL MODEL OF TRMS TRMS, shown in Fig. 1, consists of pitch and yaw rotor which are led by two DC motors and placed on the both ends of a beam. TRMS can vertically and horizontally rotate in pitch and yaw axis respectively. The control goal is to control the position of pitch angle \u03c8 and the position of the yaw angle \u03d5 . The momentum equations of pitch movement are given below [1]: GBFG11 MMMMI \u2212\u2212\u2212=\u03c8 \u03c8 (1) 11 2 111 baM \u03c4+\u03c4= (2) )sin(MM gFG \u03c8= (3) )(signBBM 21B \u03c8+\u03c8= \u03c8\u03c8\u03c8  (4) )cos(MKM 1gyG \u03c8\u03d5=  (5) 1011 1 1 1 1m TsT k u G + =\u03c4= (6) Here, 22 1 m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.7-1.png",
        "caption": "Fig. 34.7 Structural analysis of semi-span wing",
        "texts": [
            " Fully assembled aircraft has the overall dimensions as listed in Table 34.10. Considering the wing as a cantilever beam a structural analysis was performed by adding a load of 120 N (calculated using ultimate load factor) on the tip of the wing and fixing the root of the wing showed that the maximum stress was on the carbon fibre spar. The maximum stress on the carbon fibre spar is 525.28 MPa, whereas the ultimate tensile stress of carbon fibre is 1500 MPa. Hence, the wing has a factor of safety of 2.85 (Fig. 34.7). The enclosure size was taken as 70 * 20 * 20m in x-, y- and z-directions, respectively. Sufficient care was taken to make sure the mesh quality is good. The skewness of the wing was 0.85. A sizing of 10 mm was given to the wing, and an inflation layer with a y + value of 1 with a growth rate of 20% was added to capture the boundary layer. Mesh independence test revealed that this setting is a perfect trade-off between computation time and accuracy. We chose a pressure-based steady-state simulation, and the aircraft was analyzed at cruise velocity using Spalart\u2013Allmaras model [11]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure11-1.png",
        "caption": "Fig. 11: Model reduction",
        "texts": [
            " In a 3D-approach, the nodes can be deflected in three degrees of freedom (x-, y- and z-direction). For this reason, a modal matrix \u03a8\u20320 = [~\u03a8\u20320,1, ~\u03a8 \u2032 0,2, ..., ~\u03a8 \u2032 0,i] (11) for each degree of freedom and each of the i modes is extracted. Considering a 3D-model creates large matrices and causes a high computational effort. In the following steps, only the transfer function of the forces acting on the inner stator to the vibration at the outer housing surface is required. Therefore, all nodes inside the structure are irrelevant and are neglected. Fig. 11 exhibits on the left the complete model with all nodes. On the right, the reduced model is shown as it is adopted for the use in Matlab. The red points represent the nodes on which the forces act, the blue points mark the surface on which the deflection is analyzed. For a 2D-FFT performed later on the outer surface, the nodes and eigenvectors are interpolated to equidistant coordinates in r- and \u03b3\u2032-direction. A prerequisite for this approach is a relatively cylindrical shape, as it is mostly found for water jacket cooled traction motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001102_j.matpr.2020.05.310-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001102_j.matpr.2020.05.310-Figure2-1.png",
        "caption": "Fig. 2. Meshed view of the wheel housing of a four wheeler.",
        "texts": [
            " But real structures cannot often be modeled using such simple elements so alternative methods are sought to develop [K] when shape, shape, material properties and behavior are more complex. Both Vibrational and weighted residual methods have been developed to derive [K] using piecewise continuous distribution for the nodal unknowns. For structural mechanics problems it is observed that the finite element characteristics are identical whichever of the two methods is used. The wheel housing designed was meshed in the software for proceeding with the nodal analysis to be performed in the ANSYS software [23]. The meshed view is shown in the Fig. 2. The mesh is then opened in the software for the analysis in which initially the load is applied to check for the stresses. The load applied gets enforced in the housing which is portrayed in the Fig. 3. On applying the stresses in the housing the deflection is noted in the output as 0.42 mm for the applied load as shown in the Fig. 4. The stress acting on the housing of the wheel while applying load is 39.008 N/mm2 as al., FEA based approach on replacing the metal cast wheel into thermoset 20.05"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001074_s12046-020-01389-z-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001074_s12046-020-01389-z-Figure5-1.png",
        "caption": "Figure 5. Basic structure (a) Stator with no notch, (b) Stator with notch.",
        "texts": [
            " This paper analyses all the designed motors. However, since proposed LSRM contains a stator design modification due to addition of a notch therefore the designed dimensions of the 6/16 LSRM are further corrected with parameter based cumulative deterministic optimization algorithm (PBCDOA) [13]. To start the optimization process, initially the LSRM is designed with 16 translator teeth having no stator notch. However, the breadth of the stator teeth was increased to maintain the number of stator phases constant. Figure 5 shows the structural difference between the LSRM with zero stator notch and LSRM with notch. For further reference, LSRM without notch will be referred as LSRM (0) and un-optimized LSRM with notch will be referred as LSRM (1). Fill-factor of the slot, turns count/phase, heights of stator and translator, pole-pitches of stator and translator, air-gap, and stack length are maintained constant during the entire optimization. Figure 6 shows the algorithm used for the optimization process. Table 2 shows force performance comparison of the conventional LSRMs and LSRM (0)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure20-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure20-1.png",
        "caption": "Fig. 20 Working configuration of the metamorphic nipper swing mechanism",
        "texts": [
            " Taking the metamorphic nipper swing mechanism (Zhang and Sun 2015) as an example, the types and constraints of metamorphic joints based on the kinematic characteristics of metamorphic joints are given. The mechanism has three configurations, which are: the nipper is gradually closed, the nipper is closed, and the nipper is gradually opened during working cycle. Based on that, the nipper gradually closed configuration and the nipper gradually opened configuration are taken as a mechanism in order to achieve the above requirements, as shown in Fig.\u00a020a. In this mechanism, spring is added to kinematic joint G, so that the slider 5 and the component 6 are combined into one component under the constraint of spring, and the nipper is in an open state at this time. The nipper closed configuration is shown in Fig.\u00a020b. In this mechanism, a geometric constraint is added at kinematic joint C, and the upper nipper 4 which combined with the lower nipper 2 to form one component is gradually moved from position b to position a. Based on the principle of augmented Assur groups, the metamorphic nipper swing mechanism is divided into a fixed-axis rotating active part, an Assur group RPR and an augmented Assur group RP\u2013RR\u2013RR\u2013R, as shown in Fig.\u00a021. According to geometric and inertia properties of the metamorphic nipper swing mechanism in Table\u00a0 3, a three-dimensional model is established in SolidWorks, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001835_j.precisioneng.2020.11.006-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001835_j.precisioneng.2020.11.006-Figure3-1.png",
        "caption": "Fig. 3. Schematics of mechanical drive module.",
        "texts": [
            " Particularly, the input data of deep CNN model is divided into a set of segments in this study. This data preprocessing method is to jointly express the temporal correlations in time sequence and the coupling relationships among the factors. The model of the feed drive is constructed as shown in Fig. 2, which includes feed-forward function module, servo control module and mechanical drive module. This model was built in our previous work,[30] a brief introduction is made in this part. The physical structure of the ball screw feed drive module is shown in Fig. 3. It mainly consists of three parts: motor drive, the ball screw drive and the mechanical execution. The symbols in Fig. 3 are illustrated in Table 1. The corresponding mathematical expression of ball screw feed drive module is as follows: T = Jm\u03b8\u0308m + cb\u03b8\u0307m + fb + l 2\u03c0 F (1) Table 1 Definitions of symbols. symbols Definitions symbols definitions Tm Torque of motor xt Displacement of axis \u03b8m Rotation angle of motor Mt Equivalent mass of workbench Jm Moment of inertia of motor c\u2217 Viscous damping K Equivalent axial stiffness f\u2217 Friction L Screw lead J. Zhang et al. Precision Engineering 68 (2021) 82\u201396 F =K ( l 2\u03c0\u03b8m \u2212 xt ) + cb ( l 2\u03c0\u03b8\u0307m \u2212 \u03b8\u0307t ) (3) The relationship between the output angle of the motor and the displacement of the actuator is obtained as: l 2\u03c0 ( K\u03b8m + cb\u03b8\u0307m ) =Mtx\u0308t +(ct + cb)x\u0307t +Kxt + ft (4) When the servo control module is added to the mechanical drive model shown in Fig. 3, a common ball screw feed drive model is obtained as shown in Fig. 4. J. Zhang et al. Precision Engineering 68 (2021) 82\u201396 By referring to the control theory,[31] the transfer functions of feed drives can be simplified as follow xt(t) = xi(t) + \u03d5v.x \u2032 i(t) + \u03d5a.x \u2032\u2032 i (t) (5) where \u23a7 \u23aa \u23aa\u23a8 \u23aa \u23aa\u23a9 \u03d5v = 2\u03c0 Kpp.l \u03d5a = \u2212 2\u03c0 Kpp.l \u00d7 [ \u2212 2\u03c0 Kpp.l + ct K + Ti DA.Trg.Kvi \u00d7 ( cb + ct. l2 4\u03c02 )] (6) To reduce the tracking error of the feed drive system, the signal is compensated by a feed-forward function module, and the compensated input xie is as follow: xie = xi + Kpf "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002004_ddcls49620.2020.9275203-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002004_ddcls49620.2020.9275203-Figure3-1.png",
        "caption": "Fig. 3: Simplified vehicle Model",
        "texts": [
            " Consequently, transformation from body to right wing frame is calculated using a combination of the rotation matrices as given in (7). RWR,SR = R(\u03bb)R(\u03d5) (6) RWRB = RWRSRRSRB (7) Since the tail has a degree of freedom with respect to the body, in order to describe the swing of the tail, it is necessary to define a tail reference frame, which is similar to the B frame, but with the x \u2212 axis and the y \u2212 axis in opposite directions. The rotation matrix from the tail frame to B frame can be written as (8). Finally, the simplified aircraft model and reference frames are shown in Figure 3. RTB =  cos(\u03b8T ) 0 \u2212sin(\u03b8T ) 0 1 0 sin(\u03b8T ) 0 cos(\u03b8T )  \u22121 0 0 0 \u22121 0 0 0 1  (8) The movement of the vehicle can be divided into translation and rotation. Here, both nonlinear equations of motion are presented in Lagrangian form. In order to control the position and attitude of the flapping wing aerial vehicle, the following vector is considered as the state variables: q = [qt, qr] T = [x, y, z, \u03c6, \u03b8, \u03c8]T (9) where x, y and z are the position of the vehicle in the inertial frame, \u03c6, \u03b8 and \u03c8 are Euler angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002423_012128-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002423_012128-Figure2-1.png",
        "caption": "Figure 2. Repeating unit. Figure 3. Prototype of unit cell.",
        "texts": [
            " Designing an auxetic structure that has a simple geometry with a unit cell that provides room to be easily scaled without compromising its auxetic nature is a challenging task. This paper has addressed this my modifying the idealised unit cell derived from an auxetic foam [5] to make it more suitable for mass production. The auxetic structure is selected and modified based on four key factors. The structure should have less repeating units, better scalability, use standardised raw materials and mainly the feasibility for mass production. A new re-entrant cubic auxetic structure Figure 1 is developed based on these key factors. A repeating unit shown in Figure 2 is a sub unit that constitute a unit cell (Figure 3). This unit cell exhibits auxetic properties. The lesser the repeating units the easier would be the assembly of the unit cell. These unit cells must have the provision to link with other such unit cells so it can be scaled with minimal effort. The materials used to fabricate the auxetic structure must be available in standard dimensions. This eliminates the additional step of in house fabrication. Finally the shape of the structure must be simple and cost effective to be assembled. The cubic unit cell (Figure 1) that resembles a systematically collapsed cube is re-entrant in nature. This cell can be divided into repeating units that can be further subdivided into standard geometric shapes. Based on the availability of standard sized raw materials aluminium cubes (6.1 mm) and steel links (\u00d8 1.3 mm) are selected. These are fixed parameters in the unit cell. The repeating unit (Figure 2) consists of 5 steel links assembled in a particular formation to facilitate assembly with other repeating units. The re-entrant cubic unit is an assembly of 6 repeating units (Figure 4). The dimensions of the unit cell assembly can be controlled with two parameters, length of the steel link (\ud835\udc59) and angle (\ud835\udf03) between the steel links (Figure 2). ICRIET 2020 IOP Conf. Series: Materials Science and Engineering 1070 (2021) 012128 IOP Publishing doi:10.1088/1757-899X/1070/1/012128 These parameters are assigned 3 levels to determine the optimum combination (Table 1). The desired functionality of the unit cell is to provide the maximum negative Poisson\u2019s ratio with minimal induced stress, while subjected to an external force. L9 orthogonal array is used to analyse feasible test cases (Table 2). A compressive load of 25 N is applied in the axial direction of the cubic unit cell with its bottom link fixed (Figure 5)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure9-1.png",
        "caption": "FIGURE 9. The von Misses stress of modified frame",
        "texts": [
            " According to result in the Figure 5 and Figure 6, the frame must be modified to enhance its performance since the structure can be considered as fail. The value of intrusion and also the stress were not fulfilling the requirement of ECE 66 standard. Some part of the frame has been modified by increasing its thickness to reduce the stress and also the value of intrusion. List of modification was tabulated in Table 3. The location of modified part in the frame were depicted on Figure 7. 030153-5 The simulation was done for the modified frame and the result for displacement, stress and also intrusion were depicted in Figure 8, Figure 9 and Figure 10 respectively. Based on the result of simulation, the displacement of frame was decrease. The magnitude of maximum deflection is 359.4 mm, and minimum displacement is 39.94 mm. 030153-6 The von Misses stress actually increase compared to the existing frame, but during the rolling accident, the collapse part is needed to absorb the impact energy from the external load. Therefore, the passenger is safe since they received the minimum impact energy. The maximum von Misses stress is 230"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002124_marss49294.2020.9307898-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002124_marss49294.2020.9307898-Figure1-1.png",
        "caption": "Fig. 1. The concept diagram of non-invasive microrobot system. The wireless magnetic manipulation system for magnetic microrobot will be incorporated with the integrated ultrasound and photoacoustic (USPA) imaging system for complete non-invasive tracking and operation of future intelligent microrobot. The right-hand figure shows a optical microscopic image of four different sizes of microrobot laying around a human hair.",
        "texts": [
            " PA imaging is molecule sensitive since the ultrasonic signal is generated due to optical absorption. As such, the microrobot agent made from magnetic metallic material is then able to significantly boost the imaging contrast of PA method. Therefore, this integrated USPA method is more sensitive than ultrasound alone and can provide more functional and molecular information [23]\u2013[26]. Therefore, the work presented in this paper is our initial effort to incorporate integrated USPA imaging in order to track single microrobot in non-transparent environment (Fig. 1), which will allow for the whole microrobot system working in non-invasive manner. The following section II introduces the magnetic microrobot prototypes used in this study and the experimental setup of USPA imaging. The experimental results of the imaging tests will be demonstrated and analyzed in section III. In the conclusion section IV, further integration of the microrobot and photoacoustic imaging will be discussed in order for potential real biomedical operation trials. Series of magnetic microrobots are in different sizes prepared for the imaging tests using integrated USPA method",
            " The average diameter of the particles is approximately 3 \u2212 7 \u00b5m that fits in the microrobot\u2019s envelope dimension well. The 99.9% metals basis particle and also its dark grey to black color will induce strong absorption of photonic signals, which generates strong imaging contrast. The Ni particles were mixed in the negative photoresist (SU-8 2035, MicroChem) in a density of 1.25 g/mL (Volume ratio \u2248 1 : 7). The sample was then patterned through photolithography to produce the microrobot prototypes in varied dimensions (400 \u00b5m, 200 \u00b5m, 100 \u00b5m, 50 \u00b5m square, Fig. 1). The thickness of the developed micororobots is approximately 40 \u00b5m. The varied sizes of single microrobot agent will not only gauge the imaging sensitivity of the USPA modality, but also compare out the capacity of USPA and pure US method. B. Integrated Ultrasound and Photoacoustic (USPA) Imaging The integrated USPA is an advanced biomedical imaging method combining the advantages from both US and PA. It provides co-registered bio-tissue background morphology (US) with a high contrast image of the microrobot (PA) at relatively high penetration depth [27]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002374_j.matpr.2020.11.841-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002374_j.matpr.2020.11.841-Figure5-1.png",
        "caption": "Fig. 5. Von moises stress and dis",
        "texts": [
            " Vibration test fixtures are required to allow mounting of the test specimen to the fix as well as to allow for testing in all the three orthogonal directions. The design of vibration test fixtures is critical to avoid errors in equipment test response due to any resonances of the impact hammer, table and the fixture itself [12]. Ideally the laboratory mounting should replicate the physical conditions observed in service such as the stiffness, mass and the consequent resonant responses of the actual service installation. The model of car seat and the deformation car seat is shown in Fig. 4. Fig. 5 shows the Von moises stress and displacement in the car sheet. The deformation present in car sheet surface with frequency analysis is shown in Fig. 6. The details of stress versus displacement of ABS, polyurethane and Nylon materials observed by Modal (without bush) analysis are given in Tables 1\u20133. Cost reduction. Less waste. Reduce production time. An enhanced competitive advantage. Reduce errors. Confidentiality. Production on demand. Surface texture is generally too rough. Materials have low heat deflection temperatures Materials generally have low strengths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001707_0309324720958257-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001707_0309324720958257-Figure8-1.png",
        "caption": "Figure 8. Forced analysis of braided wire rope strands.",
        "texts": [
            " For example, the right-hand braided strand 1 with an initial phase angle of 0 is selected, and the unfolded line of the strand center line is as shown in Figure 7(b). Each strand in the braided wire rope is treated as a whole. Considering the unfolding line of the jth segment of the ith strand as an example, the axial tension along the strand is Fij, shear force is Qij, component of the two forces in the axis direction of the braided wire rope is Tij, component of the horizontal direction is Pij, torque of the strand is Mij, and braiding angle after loading is b\u2019ij that is defined as the center line of the braided wire rope strand. As shown in Figure 8, i is the ith strand of the braided wire rope, i=1, 2, 3, ., 8; j is the jth strand of a pitch within a pitch Segment, j=1, 2, 3, 4, 5. The following equations can be obtained: Tij =Fij cosb 0 ij +Qij sinb 0 ij Pij =Fij sinb 0 ij +Qij cosb 0 ij ( \u00f010\u00de According to the literature,17 the strand shear force Qij is calculated as follows: Qij = EgpR4 g 4 sin2b 0 ij r 0 ij + sin2bij rij (eij 1) ! sinb 0 ij cosb 0 ij r 0 ij + EgpR4 g 4(1+ n) sinb 0 ij cosb 0 ij r 0 ij + sinbij cosbij rij (eij 1) ! sin2b 0 ij r 0 ij \u00f011\u00de The braided angle before and after deformation has the following relationship: sinb 0 ij = 1+ nisin 2bij jijnicos 2bij 1+ nisin 2bij + jijcos 2bij sinbij cosb 0 ij = (1+ jij)(1+ nisin 2bij) 1+ nisin 2bij + jijcos 2bij cosbij 8>>< >>>: \u00f012\u00de In the above equation, ni = Dxij=xij DSij=Sij is the Poisson\u2019s ratio, jij= Dlij lij is the axial strain along the rope of the jth segment of the ith strand, lij is the length of the wire rope corresponding the jth segment of the strand i, Dlij is the deformation amount in the wire rope after the load, Sij is the length of the strand, DSij is the deformation amount in the strand after the load, xij is the corresponding elliptical arc length, and Dxij is the variation in the arc length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001993_ddcls49620.2020.9275125-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001993_ddcls49620.2020.9275125-Figure6-1.png",
        "caption": "Fig. 6: Collision detection between manipulator and obstacle Let I = \u2016- \u2212 - 7!\u2016 , A = K- E \u2212 - 7!K , D = K- E \u2212- K. When I\" ; A\" L D\", as the situation in Figure 6 (a), the closest distance between the link and the obstacle is A \u2212F E \u2212 F G; when I\" ; D\" L A\", as the situation in Figure 6 (b), The closest distance between the link and the obstacle is D \u2212 F E \u2212 F G ; when a, b, and c satisfy the situation in Figure 6 (c), the closest distance is - EH \u2212 F E \u2212 F G , and the value of - EH can be calculated by the following formula : - EH = 2N/I (4)",
        "texts": [
            " Therefore, in this paper we will sample nodes in the joint space, and the collision detection process will be performed in Cartesian space. The pose of each link corresponding to each node in joint space are calculated through forward kinematics. The positional relationship between the link and the obstacle is used for collision detection. In this paper we use a collision detection method based on sphere bounding box. Place the obstacle in a properly sized sphere that completely surrounds the obstacle. When collision detection, the manipulator link can be abstracted into a cylinder model. Figure 6 shows the three relative positions of a link and the obstacle, where > ? represents the cylinder model, @ABCD represents the sphere bounding box, - is the position of the joint , - E is the position of the obstacle, F G is the radius of the link, F E is the radius of the obstacle, and H is the foot from the center of the ball to the link. where S is the area of triangle - E- 7!- , which can be calculated by Helen's formula: N = PQ(Q \u2212 I)(Q \u2212 A)(Q \u2212 D) (5) where Q = (I ; A ; D)/2. So the distance between the link and the obstacle is R = S A \u2212 F E \u2212 F G ,D \u2212 F E \u2212 F G ,- EH \u2212 F E \u2212 F G , I\" ; A\" L D\"I\" ; D\" L A\"T \u210eCF$ (6) When R L 0, the obstacle is considered to collide with the link"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure52.4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure52.4-1.png",
        "caption": "Fig. 52.4 a Rotary cutter unit; b support frame unit; c arm locking arc; d lifting unit",
        "texts": [
            " The power required for cutting the soil to the rotary cutting unit will be provided by the engine mounted on top of the rotary cutting unit. Each individual partwere created first in \u201cPartDesign\u201dworkbench and later assembled on \u201cAssembly Design\u201d workbench. The assembled design of the system was simulated on \u201cDMU Kinematics\u201d workbench (Fig. 52.3). It consists of six blades made of tool steel of 5 mm thickness mounted on three flanges mounted on the shaft casing pipe. The flanges are in the shape of a circular disc having six cut out slots along the circumference in order to house the blades (Fig. 52.4a). The support frame unit consists of the main frame upon which the engine will be mounted. The bearings and the shafts to be connected to the rotary cutting unit are mounted to this frame as well. A plate in the shape of an arc which provides locking positions of the arm, and the support frame is also fixed to this frame. A positioning pin is fixed to the frame to act as the pivot point about which the frame will rotate with respect to the arm. The support frame with its constituent components is shown in Fig. 52.4b. The positioning plate has a number of slots on it where the position of the support frame can be fixedwith respect to the armwith the help of a pin as shown in Fig. 52.4c. The total angle about which the frame can rotate is designed to be 90\u00b0. This angle can be increased or decreased depending on the requirement. Hilly terrains present varying degrees of slope along its range and a mechanism or a system with a fixed cutting position of cutting height is not always suitable for forming terrace fields. To address this problem, a mechanism for movement along the vertical direction (lifting and lowering the rotary cutting unit) was designed. The mechanism of the lifting unit is proposed as a solution so that the rotary cutting can be lowered or lifted to a specified height. This unit consists of an arm which is jointed to the support frame and a pair of lifting arms fixed to each side of a bracket Fig. 52.4d. This unit, in turn, is connected to the power tiller by means of a shaft emerging from the power tiller. In hilly terrains manoeuvring power tillers or even vehicles for that matter is a challenge due to factors such as steep ascend or descend, poor road conditions, narrow roads, lack of motorable roads and landslides, etc. During rainy seasons, it becomes evenmore dangerous for operators of suchmachines. Therefore, a modified wheel design is proposed for the power tiller so that it can provide sufficient traction while working on such terrains"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001147_j.jmmm.2020.167119-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001147_j.jmmm.2020.167119-Figure8-1.png",
        "caption": "Fig. 8. Magnetic field distribution of polyphase IM.",
        "texts": [
            " Result and discussions The algorithm of the proposed approach, shown in Fig. 5, is utilized to obtain the magnetic field analysis, mutual inductance and equivalent circuit, and instantaneous torque. The presented results will be performed at 6 phases in the stator and 6 phases in the rotor. The winding distribution is presented in Fig. 6. The machine winding configuration for one pole is illustrated in Fig. 7. By using FEMM software the magnetic field of the polyphase IM can be deduced as shown in Fig. 8. Also, the mesh was applied by the FEMM preprocessor for breaking machine geometry into several small triangular elements. The number of obtained nodes is 145,716 and the number of elements is 291230. The air gap flux density between analytical formulation using Eq. (12) and the FEMM is shown in Fig. 9. In this figure, there is a slight difference between these results. The flux density analysis of a cross-section of the polyphase IM is illustrated in Fig. 10. It is found that the flux density is uniformed distributed around each pole"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure45.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure45.2-1.png",
        "caption": "Fig. 45.2 Semi-circular shape of a bucket for ginger seed metering unit",
        "texts": [
            " The power of the metering mechanism was obtained from the furrow wheel using suitable gears. The semi-circular seed metering cup with chain and sprocket mechanism was fabricated. The length of the chain was 1420 mm and sprocket with 15 teeth on the metering mechanism and 11 teeth on the furrow wheel to form a gear train mechanism.Further, a rodoperated by a camattached to the shaft of themeteringmechanism was used as an agitator for a smooth flow of ginger seed. The semi-circular bucket of dimension 35 \u00d7 45 \u00d7 24 mm was fabricated (Fig. 45.2). The design of a power-operated mini ginger planter was conceptualized and fabricated. The developedmachinewas attachedwith a power tillerwhich is awalk-behind type. The design constitutes hopper (1), metering bucket (2), chain (3), sprocket (4), metering device covering unit (5) frame (6), furrow opener (7), furrow covering device (8), and furrow wheel (9). The developed planter consists of two furrow wheel and a horizontal frame that supports all other components. The various components of the ginger planter were shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001044_s12206-020-0533-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001044_s12206-020-0533-5-Figure1-1.png",
        "caption": "Fig. 1. A schematic view of the considered TLRM with the counterweights.",
        "texts": [
            " The point-to-point motion is a repetitive motion between two given positions. This motion contains forward and return motions. For the forward motion, the robot moves from the start position to end position. Then, in the return motion, the robot moves from the end position to start position. The point-to-point motion uses in many applications, one which is pick-and-place operation. For implementation of the ZPB method, the optimal control method can be used to obtain necessary conditions for optimality [27]. In this paper, a robot with the counterweights shown in Fig. 1 was considered. As shown on the figure, two counterweights were attached to the links 1 and 2. Herein mci, lci and \u03b2i (i = 1, 2) are mass, length, and installation angle of the counterweight i with reference to the link i, respectively. The parameters of the TLRM shown in Fig. 1 are introduced in Table 1. The Lagrange's method was used to obtain the dynamic equation. This method is written in terms of the Lagrangian function L, the generalized coordinate q, and the generalized force or torque Q, as follows: ( ) .\u2202 \u2202\u2212 = \u2202 \u2202 d L L Q dt q q (1) By definition, function L evaluates the difference between ki- netic and potential energies of a system; i.e., L = K-U, where K is the kinetic energy and U is the potential energy of the system, both written in terms of generalized coordinates q",
            " Constrained to the 16 boundary conditions given in Eqs. (18) and (22), these equations construct a twopoint boundary value problem (TPBVP). This TPBVP is solved using the bvp4c command in MATLAB software to determine the state vector x, co-state vector \u03c8, new state vector \u03bc and parameters vector b. Also, to obtain the static balancing, the gravity terms G = [G1 G2]T must be achieved to zero. So, the counterweights parameters are obtained as follows: In this section, results of simulating the TLRM shown in Fig. 1 are presented. Two cases were considered for this purpose: Unbalanced and zero-power balanced. In the unbalanced case, counterweights were zero (mc1 = mc2 = 0). The experimental set-up presented in Sec. 3.2 is based on the robot parameter values given in Table 2. Simulations and experiments were performed for different point-to-point motions. For both unbalanced and ZPB cases, the theoretical results were compared to experimental data. Consider a repetitive point-to-point motion with the following boundary conditions for the forward motion, 1 1 2 2 3 3 4 4 (0) 210 , ( ) 310 , (0) 15 , ( ) 60 (0) ( ) (0) ( ) 0, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001884_1350650120975519-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001884_1350650120975519-Figure3-1.png",
        "caption": "Figure 3. Beam on elastic foundation. (a) Whole dabber. (b) Contact part. (c) Cross section.",
        "texts": [
            " It is noted that the index n physically describes and determines the uniformity of contact pressure distribution. The greater n is, the more concentrated the distribution, while the smaller n is, the more dispersed the distribution. Differently, the distribution pattern of axial contact pressure cannot be entirely explained by the previous complete spherical contact theory due to the limited axial width of the bearing. Thus, it is necessary to analyze quantitatively the deformation of the bearing to describe the axial contact stress on the contact surface. As shown in Figure 3(a), the dabber supported on the inner ring is simplified as a circular beam lying on elastic foundation with uniform external loads P applied in the plane of symmetry of the cross section on both sides with the length of L B 2 where L is the total length of the dabber and B is the axial width of the outer ring. According to Winkler\u2019s assumption in 1867,28 the reaction forces R\u00f0z\u00de of the foundation are proportional at every point to the deflection y\u00f0z\u00de of the beam at this point: R z\u00f0 \u00de \u00bc k y\u00f0z\u00de, where k is the constant of the foundation, which includes the effect of the width of the beam",
            " It is assumed that the axis can be regarded as a one-dimensional bar with a unified width of d equal to the diameter of the circular section, and the thickness t of the inner ring keeps the same in the contact area. Then the contact k takes the form of Ed t , where E is the Young\u2019s modulus of the bearing material. The equivalent bending moment M and shear force Q on the two ends of the contact part with the length of B are then calculated to analyze the deflection of the axis, where M equals to L B 2 2 Pd and Q equals to L B 2 Pd, as depicted in Figure 3(b). The related geometrical sizes are also marked on the cross section of the structure in Figure 3(c). The analytical solutions to the problem that a straight beam supported along its entire length by an elastic medium in Figure 3(b) are where y\u00f0z\u00de, u\u00f0z\u00de, y0 and u0 stand for the deflection and the slope of the beam and their values at the original point where z \u00bc 0, respectively;M z\u00f0 \u00de and Q\u00f0z\u00de represents the bending moment and the vertical force subjected on the beam respectively; the parameter a equals to ffiffiffiffiffi k 4EI 4 q ; where the inertial moment I of the cross-section is pd4 64 . The functions A1 az\u00f0 \u00de; A2 az\u00f0 \u00de;A3 az\u00f0 \u00de;A4 az\u00f0 \u00de in equation (2) are introduced in the following way A1 az\u00f0 \u00de \u00bc ch az cos az (3) A2 az\u00f0 \u00de \u00bc ch az sinaz\u00fe sh az cos az 2 (4) A3 az\u00f0 \u00de \u00bc sh az sinaz 2 (5) A4 az\u00f0 \u00de \u00bc \u00f0ch az sinaz sh az cos az\u00de 4 (6) It should be noted that the values of y\u00f0z\u00de and u\u00f0z\u00de at the point z \u00bc B are equal to y0 and u0 respectively according to Figure 3(b). Thus the values of y0 and u0 can be determined by substituting z \u00bc B into equation (2). As mentioned before, it is hypothesized that the thickness t of the foundation remains a constant all over the contact area. However, since the role of elastic foundation is played by the inner ring of the bearing, the value of t actually changes with the axial width z in a way as follows t z\u00f0 \u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi dk 2 2 \u00f0z B=2\u00de2 s d 2 0 z B (7) Here the value of t is set as the average thickness t of the inner ring over the length of B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001747_physrevapplied.14.044026-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001747_physrevapplied.14.044026-Figure6-1.png",
        "caption": "FIG. 6. Schematic representation of the double spring-mass model. Two masses located at zt and zb interact with one spring of stiffness kt. The bottom mass is connected to the moving plate with a spring of stiffness kb. The plate position is zp .",
        "texts": [
            " Actually, the lower the mass of the projectile, the lower the mass of the throwing instrument should be, in order to achieve the best efficiency. If the mass cannot be adjusted, projectiles designed for elastic-assisted throwing seem to be a very good option. We thank Fabrice Mortessagne for his support of the project, and acknowledge the financial support of CNRS. F.C. and C.R. thank Lorenzo Betti and Nicolas David for their input at the beginning of the project. We model each part of the bilayered projectile with a spring-mass system, as depicted in Fig. 6. We approximate the upper part of the projectile with a point mass mt, located at z = zt, and the bottom part with a point mass mb, which is located at z = zb. These two components are connected together with a spring of stiffness kt and length lt. The engine is modeled through a plate, located at z = zp(t), which triggers the bottom mass through an elastic spring of stiffness kb and length lb. The dynamics of this model is described by mtz\u0308t = kt(zt \u2212 zb \u2212 lt), mbz\u0308b = \u2212kt(zt \u2212 zb \u2212 lt) \u2212 kb(zb \u2212 zp \u2212 lb), zp = A[1 \u2212 cos(\u03c9t)]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002938_012002-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002938_012002-Figure1-1.png",
        "caption": "Figure 1. (a) The initial model, (b) the simplified model, (c) the model of traditional mold, (d) the model of conformal cooling mold, (e) the model of traditional channels, (f) the model of conformal cooling channels",
        "texts": [
            " Fluent is MEMAT 2021 Journal of Physics: Conference Series 1939 (2021) 012002 IOP Publishing doi:10.1088/1742-6596/1939/1/012002 used to simulate the cooling effect of traditional cooling channels and conformal cooling channles. It will analyze the advantages of conformal cooling channels from the aspects of cooling effectiveness, efficiency and uniformity, and the die-casting mold insert is manufactured by the SLM equipment. The conformal cooling channels is designed for the die-casting mold insert. The initial model of the mold insert is showed in Fig. 1(a). The mold surface in the insert that needs to be cooled is showed in Fig. 1(a). In order to facilitate the design of the conformal cooling channels and reduce the calculation amount of the heat flow simulation, the simualtion mold is simplified as shown is Fig. 1(b). The traditional cooling channels of the die-casting mold insert is formed by drilling, shown in Fig. 1(c) and Fig. 1(e). The conformal cooling channels model is shown in Fig. 1(d) and Fig. 1(f). Fluent is used to simulate the cooling process of the mold insert. Fluent is a commercial CFD software, which is widely used in the aerospace, automotive, medical and electronics industries. Fluent\u2019s calculations are based on the finite volume method. It targets the control volume rather than the mesh modes and discretizes the integrals of conservation equations in physical space. Finite volume method can be adapted to structured and unstructured grids, and its general approximation accuracy is second order"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000144_elecsym.2019.8901645-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000144_elecsym.2019.8901645-Figure4-1.png",
        "caption": "Fig. 4. Three omni directional robot representation",
        "texts": [
            " Position and orientation affect the direction of the robot. Mobile robot in this research equipped with a three omni directional wheel. In order to determine position, mobile robot has three rotary encoder sensor attached to each of its dc motor. By utilizing rotary encoder sensor that is attached to mobile robot\u2019s wheel, the position of mobile robot can be discovered using forward kinematics of three wheeled omni directional robot. Representation of three omni directional wheeled robot as shown in fig. 4. Based on kinematic representation of three omni directional wheeled robot, velocity of the robot in x and y axis can be calculated by knowing velocity of each omni directional wheel. Equation (3) and (4) describe the formula to calculate mobile robot\u2019s velocity on x and y axis. Vx = V3 \u2013 V1 Cos(\u03b8) \u2013 V2 Cos(\u03b8) (3) Vy = V1 Sin(\u03b8) \u2013 V2 Sin(\u03b8) (4) After mobile robot\u2019s velocity on x (Vx) and y axis (Vy) known then distance traveled by mobile robot can be calculated. If mobile robot\u2019s velocity is known then the next step is to find its position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure11-1.png",
        "caption": "Fig. 11. (a) External contact and (b) internal contact between two cylinders.",
        "texts": [
            " It should be noted that to evaluate the forces accurately, contact force formulations that consider physical phenomena and wide range of applicability are required. The next sub-section elaborates on different contact mechanics formulations and their applicability. The contact between rotor and roller is modelled as contact between two cylinders of equivalent radii of curvature (detailed in Section 3.3.2). To evaluate the contact force, a suitable cylindrical contact force model, applicable to both external and internal contact (Fig. 11) is required. Hertz [17] pioneered the study of contact forces between two elastic bodies and concluded that the contact area was, in general, elliptical. Hertz\u2019s law relates the contact force with a nonlinear power function of indentation and can be expressed as: FN \u00bc Kdn \u00f010\u00de where d represents the relative indentation between the contacting bodies, K and n are the contact stiffness parameter and the nonlinear power exponent determined from material and geometric properties of the local region of the contact. Limi- tations such as applicability to conformal contacts (typically observed in internal cylindrical contacts, Fig. 11(b)), evaluation of K;n, and assumption of elastic contact (not considering the energy dissipation during the contact process) limit its applicability. Johnson [18] provided a contact model applicable to cylindrical bodies with parallel axes in terms of their radii, Ri and Rj and axial length b as: FN \u00bc pLE ln 4pbE DR FN 1 1 d \u00f011\u00de where DR is the radial clearance between two contacting bodies, DR \u00bc Ri Rj for internal contact and DR \u00bc Ri \u00fe Rj for external contact. E is the Effective Young\u2019s Modulus, which can be evaluated for two contacting bodies with Young\u2019s moduli E1; E2 and Poisson\u2019s ratio m1; m2 as: 1 E \u00bc 1 m21 E1 \u00fe 1 m22 E2 \u00f012\u00de Johnson\u2019s cylindrical contact model, although experimentally validated in its development, has a limitation because of the logarithmic form of the model, thus limiting its applicability based on the radial clearance available and the material properties [19]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001618_ijcnn48605.2020.9207539-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001618_ijcnn48605.2020.9207539-Figure1-1.png",
        "caption": "Fig. 1: A dynamic formation learning problem for swarm-onswarm guidance based on shepherding mode of control. Each UAV is assigned an operating angle. The subgoal of the UAV is defined by angle \u03b8 and the distance \u03b4.",
        "texts": [
            " The pseudo code of the PSO-w algorithm is shown in Algorithm 1. In this paper, a PSO-based dynamic formation learning method is proposed for optimizing the formation of multiple UAVs to guide the sheep swarm towards a target position. The formation of the UAVs is determined by a set of subgoal points whose positions are generated based on the sheep\u2019s positions. These subgoals are produced by Perceptron-networks optimized by PSO. Each UAV is assigned a subgoal as its navigation destination to deploy the formation. Figure 1 illustrates a formation of UAVs to herd the sheep swarm, consisting of multiple subgoal points which are defined according to two factors: the angular position \u03b8 and the distance \u03b4. Five Perceptron-networks are evolved as subgoal production models to specialize in relative positions in the formation for multiple UAVs. Tactical subgoals are generated at different locations within 180 degree arc in each formation for guiding the sheep swarm. The Perceptron-networks are trained to specialize in locations for UAVs acting as left fielder (LF), left midfielder (LMF), midfielder (MF), right midfielder (RMF), and right fielder (RF) respectively",
            " Each UAV uses one of the production network to generate a subgoal consisting of angular position \u03b8 and distance \u03b4 relative to the sheep swarm based on the inputs. The agent then computes its normalized force vector F t\u03b2jsg towards that assigned position, then the movement of the UAV is guided by: F t\u03b2j = F t\u03b2jsg +We\u03b2jF t \u03b2j\u03b5 (10) where F t\u03b2j is the total force exerted on the UAV \u03b2j , F t\u03b2j\u03b5 is the movement noise of the UAV and We\u03b2j is the strength of the noise. Each UAV is assigned a \u03c0/9 (20 degree) operating angle relative to the sheep swarm at the LF, LMF, MF, RMF, or RF positions as demonstrated in Figure 1. The first output of the production network (in a range of [\u22121, 1]) is mapped to an adjusted angle in the range of [\u2212\u03c0/18, \u03c0/18]. Within the assigned operating angle, the subgoal point lies on a subgoal line that forms with the bisector, at the middle of the corresponding section, an angle \u03b8 \u2208 [\u2212\u03c0/18, \u03c0/18], where a positive \u03b8 indicates the subgoal line is rotated an angle of |\u03b8| counter-clockwise, and a negative \u03b8 indicates the subgoal line is rotated an angle of |\u03b8| clockwise. The location of the subgoal on the subgoal line has the distance in the range \u03b4 \u2208 [RS + R\u03c0\u03c0, RS + 5R\u03c0\u03c0] from the sheep global centre of mass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000897_j.matpr.2020.04.031-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000897_j.matpr.2020.04.031-Figure1-1.png",
        "caption": "Fig. 1. Geometric description and tooth",
        "texts": [
            " Next, the airgap permenace model is used to calculate the flux densities in each part of the SRM, the results are compared and validated by those obtained par finite elements method for flux densities and experimental results for SRM inductance. After, the Lavers method [8] and modified Steinmetz model [11] are used for estimation of iron losses. An infrared camera is used in test bench for location of the iron losses in each part of the SRM. The study considers a doubly salient q phase, 2 pole per phase SRM, with Ns and Nr tooth respectively in stator and rotor (q = Ns/2) (Fig. 1). Phases are powered by a unipolar periodic square waveform current characterized by: period T, angular frequency x, duty cycle Ton/T = 1/q = 2/Ns and phase shifted successively by 2 T/Ns. The current across the first phase is taken symmetric at the instant t = 0. In this case The Fourier decomposition of the current across the phase p (p = 1, 2. . .q) is given by the expressions (1) and (2) below: ip \u00bc X1 k\u00bc 1 ffiffiffi 2 p Ikcos kxt k\u00f0p 1\u00de2p q \u00f01\u00de Ik is the rms value of the kth order harmonic current: Ik \u00bc Iffiffiffi 2 p q sinc k q \u00f02\u00de where sinc is the cardinal sin function and I the constant value of the square waveform current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000703_s12206-020-0122-7-Figure25-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000703_s12206-020-0122-7-Figure25-1.png",
        "caption": "Fig. 25. Boundary condition for analysis of wheels.",
        "texts": [
            " The equivalent model, which was obtained by applying the temperature among the three equivalent models, was the most similar to the detailed model. Therefore, based on the equivalent model (SBT model), it was applied to the wheel of the commercial vehicles to be verified. Fig. 24 shows the wheel structures of the commercial vehicles subject to verification of the equivalent model. A total of ten bolts is fastened to this wheel structures. Two of the ten were fastened with detailed models, and the others were clamped to equivalent models. Subsequently the changes in fastening force were compared when loosening load was applied. Fig. 25 shows the boundary condition of tightening and loos- ening analysis in the wheel structures of commercial vehicles. When tightening analysis is performed, all degrees of freedom of an inner hole of a hub are constrained. Furthermore, the torque is applied to the detailed models and the temperature is assigned to the equivalent models. Likewise, when loosening analysis is performed, all degrees of freedom of an inner hole of a hub are constrained. The load points are the rigid element for each wheels, contacting the 40 degree part of the wheels to the ground"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002293_012061-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002293_012061-Figure3-1.png",
        "caption": "Figure 3: Airflow model of the compressor",
        "texts": [
            " Temperature along the compressor gas path and pressure on the aerodynamic surfaces of the blades was determined by flow calculation in Ansys CFX with the finite element method. CFD model of the compressor inter-blade channel was obtained by arranging the domains of each compressor stage in the axial and radial directions. To build a mesh of the compressor flow, the TurboGrid grid generator was used. Volumetric finite elements intended for CFD calculations were used. To reduce the required computing power, one blade was modelled for each compressor stage with the cyclic symmetry along the boundaries of the domain (Figure 3). The boundary conditions were set in the form of total inlet pressure, airflow at the compressor outlet, and rotational speed. An interface between stationary and rotating regions (Stage Mizing-Plane) was defined on the mating boundaries of regions belonging to different steps, which allows for the interpolation between mating grids. A satisfactory criterion for the convergence of the calculation was the value of the mean square residual at the level of 10\u22126. This convergence was achieved at 1200 - 1400 iterations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002170_s00170-020-06399-z-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002170_s00170-020-06399-z-Figure4-1.png",
        "caption": "Fig. 4 Dimensions, displacement boundary conditions, and scan pattern for the finite element SLM model",
        "texts": [
            " The study assesses the influence of assuming a simplified laser spatial profile, in addition to evaluating the effects of laser spot size and overlap, forming temperature, and increased laser power density to counteract stainhardening behavior when conducting multiple LSP forming treatments. A transient thermomechanical finite element (FE) model is used to predict the thermally induced distortion arising in the SLM process. Details of the thermal, mechanical, and laser heat source sub-models comprising the thermomechanical simulation are provided in Sections 2.1 to 2.3, respectively. The model geometry and SLM scan pattern is shown in Fig. 4. It includes a 3D linear hexahedral Lagrangian mesh consisting of coupled temperaturedisplacement elements with reduced integration (C3D8RT). A uniform 50-\u03bcm mesh size that balances computational expense and accuracy is used. The same bi-directional laser scan pattern is applied for each SLM layer. Note that since the distortion (and RS) in the build is strongly correlated to the thermal history, process parameters such as scan speed and time interval between layers are matched to those used in practice so as to avoid unrealistic heat loss or accumulation",
            " This boundary (or isotherm) was said to represent the solidus-liquidus transition zone. Although 316L is considered in this work (not Inconel 625), the volumetric heat flux due to the laser incidence assumes the same double ellipsoid parameters; this is justified in that the intent of this work is to assess the viability of correcting typical SLM distortion patterns using in situ LSP, rather than a rigorous predictive modeling of the actual distortion field itself. Following simulation with the described SLM model using the geometry and scan parameters in Fig. 4, as well as Tables 1 and 2, the predicted topography of the top surface is seen in Fig. 6a (based on 6561 upper surface nodes). It is noted that the topography exhibits a convex deflection similar to that described by Li et al. [15]. From surface roughness metric Rt (range), it is evident that a \u223c 9-\u03bcm vertical distortion exists between the highest and lowest points. Figure 6b shows the region of the top surface that falls outside an assigned 2-\u03bcm conformance limit (used in this work for demonstration purposes)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002721_tec.2021.3074818-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002721_tec.2021.3074818-Figure1-1.png",
        "caption": "Fig. 1 Surface-mounted PM machines. (a) RSPM. (b) FRPM.",
        "texts": [
            " Finally, the prototype machines are manufactured and tested to further verify the theoretical analysis and FEA predictions. Index Terms\u2014Flux-reversal, permanent magnet machine, finite element analysis, analytical model, superposition methods, cogging torque, UMP. I. INTRODUCTION ERMANENT magnet (PM) machine is one of the most widely utilized machine categories for industrial applications [1]-[2], including wind power generator, servo system, in-wheel motor and compressor, etc. It always houses PMs in rotor side, such as rotor surface-mounted PM (RSPM) machine [Fig. 1(a)]. However, in RSPM machines, retaining sleeve is usually necessary to protect PMs from centrifugal force, resulting in a more complicated manufacture and a larger effective air-gap. Hence, in order to avoid such demerits, more attentions are paid to a new stator surface-mounted PM machine, i.e., flux-reversal (FR) PM (FRPM) machine [1]-[2], which installs PMs on stator surface rather than rotor surface [Fig. 1(b)]. Generally speaking, there are two commonly used modeling methods, i.e., numerical methods (finite element analysis, FEA) and analytical methods. Nowadays, with the rapid development of computer technology, the commercial software becomes more and more convenient to assist the research work. Hence, many scholars employ the FEA to analyze the performances of FRPM machine, including the cogging torque, inductance, Xiaofeng Zhu and Guishu Zhao are with School of Electrical and Automation Engineering, Nanjing Normal University, Nanjing Jiangsu Province, 210023, China (e-mail: zhuxiaofeng8908@163",
            " Finally, some conclusions are drawn in section VI. Before the discussion, it should be noted that the impact of each pole is considered independent from the others. Generally speaking, the open-circuit air-gap flux density in FRPM machines is fundamental to other characteristics, e.g., backEMF, cogging torque, UMP, etc. Based on the magneto-motive force (MMF)-permeance model [24]-[25], the air-gap flux density B can be written as: \ud835\udc35(\ud835\udf03, \ud835\udefc) = \ud835\udc39\ud835\udc5d\ud835\udc5a(\ud835\udf03)\ud835\udeec(\ud835\udf03, \ud835\udefc) (1) where, \u03b8 is stator position and \u03b1 is the rotor position, as shown in Fig. 1, \ud835\udc39\ud835\udc5d\ud835\udc5a = \u2211\ud835\udc39\ud835\udc5b sin (\ud835\udc5b\ud835\udc41\ud835\udc60\ud835\udf03) is PM-excited MMF, \ud835\udeec = \ud835\udeec0 + \u2211\ud835\udeec\ud835\udc5acos\ud835\udc5a\ud835\udc41\ud835\udc5f(\ud835\udf03 + \ud835\udefc) is the air-gap permeance [17]. As can be seen from (1), the PM air-gap flux density (open-circuit) is related to the spatial relationship between stator and rotor, which means it can be modeled by superposition of a single stator tooth or a single rotor tooth if the flux leakage and magnetic saturation are neglected. Moreover, other characteristics can also be modeled by the superposition method based on either a single stator tooth or a single rotor tooth, including back-EMF, cogging torque, UMP, etc. as expressed: \ud835\udf11(\ud835\udf03, \ud835\udefc) = \ud835\udc53(\ud835\udf03) \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (2) where, \ud835\udf11(\ud835\udf03, \ud835\udefc) is the characteristic related to B(, ), \ud835\udc53(\ud835\udf03) is the function related to PM-excited MMF, and \ud835\udc54(\ud835\udf03, \ud835\udefc) is the function related to air-gap permeance. Setting the FRPM machine shown in Fig. 1(b) as an example, the mechanism of superposition of a single stator tooth can be described as follows. The open-circuit characteristics of the machine with \ud835\udc41\ud835\udc60 stator teeth [Fig. 2(a)] can be modeled by the superposition of those having a stator with single tooth only. It should be noted that the positions of the single stator tooth and rotor teeth, i.e., Figs. 2(b)-2(c) remain the same as Fig. 2(a). So does the rotor topology. Then, the characteristic of the machine with 1st stator tooth only [Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001671_icuas48674.2020.9213853-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001671_icuas48674.2020.9213853-Figure1-1.png",
        "caption": "Fig. 1: Quadrotor with a slung load on the plane XeZe.",
        "texts": [
            " The following assumptions are considered to obtain the dynamic model: 1) The quadrotor structure is rigid; 2) the cable is massless and rigid; 3) the cable is linked to the quadrotor center of mass; 4) a frictionless pin union connects the wire and the quadrotor; 5) a point concentrates the slung mass. It is worth mentioning that a crucial assumption to get the dynamic model is the one concerning the type of union between the cable and the quadrotor. However, in the literature, this assumption is often forgotten, except in [11], as Kane\u2019s equation includes this fact. It is easy to observe, in Figure 1, that the slung load position and the quadrotor center of gravity are related as follows rQ = rp \u2212 lp, (1) where rQ = [ xQ 0 zQ ]> is the quadrotor\u2019s center of mass position, rp = [ xp 0 zp ]> is the slung mass position both with respect to the inertial frame XeZe. Moreover, l is the wire length and p is a unitary vector that points in the direction of the link between the quadrotor and the slung mass, this is, p = s\u03b8p0 c\u03b8p  (2) with \u03b8p being the angle of the cable with respect to the inertial vertical axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002604_j.apacoust.2021.108063-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002604_j.apacoust.2021.108063-Figure14-1.png",
        "caption": "Fig. 14. Simulated mode shape and frequency: (a) the spring model and (b) the equivalent model.",
        "texts": [
            " In the spring model, the stiffness of the seal strip was held constant at 5 N/mm, which is an empirical value. For the equivalent model, the dynamic stiffness was obtained from an FE analysis at a 4 mm pre-compression, which was experimentally determined by conducting a seal strip gap test when the door was locked. The integration of the equivalent model into the trim-body model was implemented by CBUSH and CVISC elements in the Hypermesh software. The comparison of the experimental and simulated results is shown in Fig. 13(b) and Fig. 14. It is observed that the mode shapes derived from experiment and simulation are similar, and the errors of the mode frequency with the experimental value are 0.7 Hz in the spring model and 0.9 Hz in the proposed model. From the modal analysis standpoint, both the spring model and the equivalent model can be utilised to represent the seal strip. VTF is a crucial index in the exploration of vehicle body vibration characteristics, which allows the designer finding the most sensitive excitation points in the vibration transmission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000284_j.compeleceng.2019.106534-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000284_j.compeleceng.2019.106534-Figure2-1.png",
        "caption": "Fig. 2. Winding distribution. (a) Torque winding (b) Suspension winding.",
        "texts": [
            " In Section 6 , two prototypes are built, and their speed response and radial displacement are tested. In Section 7 , the conclusions based on the analytical method, simulations as well as experiments are drawn. Fig. 1 shows a two-dimensional model of the motor built by ANSYS Maxwell . The stator adopts double-layer windings and the slot type is pyriform. The rotor has cast aluminum and the slot type is parallel. The stator and rotor core are laminated by silicon steel sheets, with a lamination coefficient of 0.98. Table 1 lists the parameters of the BIM. Fig. 2 illustrates the distribution of torque winding and suspension winding in the stator. The pole-pair number of torque winding is 2 and the phase sequence is + A 1 , \u2212C 1 , + B 1 , \u2212A 1 , + C 1 , \u2212B 1 . The pole-pair number of suspension winding is 1 and the phase sequence is + A 2 , \u2212C 2 , + B 2 , \u2212A 2 , + C 2 , \u2212B 2 . Fig. 3 shows the rotor slot structure. The initial parameters are determined as H s0 = = 0.5 mm, H s2 = = 1.1 mm, H s1 = = 24.9 mm, B s0 = = 0.6 mm, B s1 = = 4.8 mm, and B s2 = = 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000613_icus48101.2019.8995977-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000613_icus48101.2019.8995977-Figure1-1.png",
        "caption": "FIGURE 1. Examples of address DAGs.",
        "texts": [
            " \u2022 Content XID (CID): CID is defined as the hash of a content, which allows clients to retrieve a desired content from anywhere and check its correctness. \u2022 Network XID (NID): NID specifies a network (i.e., autonomous domain) and allows clients to validate the intended network, with which it is communicating. The XIA introduces Directed Acyclic Graphs (DAGs) to represent the addresses of XIDs to achieve flexible addressing. The DAGs are highly flexible by allowing packets to express fallbacks and scoping to realize user intent. An example routing technique is given in Fig.1. A \u2018\u2018dummy\u2019\u2019 source \u2018\u2018\u2022\u2019\u2019 represents the conceptual source of a packet and the target is as a sink. Scoping means that the packet must be first routed to a scoping XID before being sent to the destination even if it has the direct route. For example in Fig.1(b), outers deliver the packet to NID first and then forward the packet to an intended SID. Another example of fallback is presented in Fig.1(c). If the SID is not available or recognized, the router will use the fallback NID through a fallback edge (dotted line). Finally, the combined mechanism of scoping and fallback in the DAG is presented in Fig.1(d). Each router along the fallback path can route directly to the intended node. Owing to the supporting CID, the XIA enables the opportunistic caching of contents for future services. Similar to the caching in the NDN, routers can also cache the data packets they receive opportunistically. Thus, data consumers can require contents by directly expressing their requests with the CID and retrieve data from either a data producer or content caches of routers. This mechanism can improve the efficiency of data retrieve and save bandwidth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001706_s12555-020-0031-7-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001706_s12555-020-0031-7-Figure4-1.png",
        "caption": "Fig. 4. Interactive torque at the second joint of a 2R planar manipulator.",
        "texts": [
            "( \u03c41 \u03c42 ) =  m2l2 2 +2m2l1l2c2 +(m1 +m2)l2 1 m2l2 2 +m2l1l2c2 m2l2 2 +m2l1l2c2 m2l2 2 (\u03b8\u03081 \u03b8\u03082 ) + ( \u2212m2l1l2s2\u03b8\u0307 2 2 \u22122m2l1l2s2\u03b8\u03071\u03b8\u03072 m2l1l2s2\u03b8\u0307 2 1 ) + ( m2l2gc12 +(m1 +m2)l1gc1 m2l2gc12 ) . (3) Equation (3) can be written as \u03c41 =\u03c4\u03032 + \u03c4\u03031; \u03c42 = \u03c4\u03032, (4) where \u03c4\u03031 =(m2l1l2c2 +(m1 +m2)l2 1)\u03b8\u03081 +m2l1l2c2\u03b8\u03082 \u2212m2l1l2s2\u03b8\u0307 2 1 \u2212m2l1l2s2\u03b8\u0307 2 2 \u22122m2l1l2s2\u03b8\u03071\u03b8\u03072 +(m1 +m2)l1gc1, (5) \u03c4\u03032 =(m2l2 2 +m2l1l2c2)\u03b8\u03081 +m2l2 2 \u03b8\u03082 +m2l1l2s2\u03b8\u0307 2 1 +m2l2gc12. (6) We may write this as \u03c4\u0303 = M\u0303(\u03b8))\u03b8\u0308 +V\u0303 (\u03b8 , \u03b8\u0307)+ G\u0303(\u03b8). (7) 3.1. Physical significance of \u03c4\u0303i: A free-body diagram illustrating the interaction between the link 1 and 2 through the second joint is shown in Fig. 4. Here \u03c4\u03032 is the net moment exerted on link 2 at the second joint. Let \u03c41 be the applied torque at the first joint on link 1. Now the net moment on link 1 is \u03c4\u03031 = \u03c41\u2212 \u03c4\u03032, or \u03c41 = \u03c4\u03031 + \u03c4\u03032. (8) Equation (8) is nothing but the first equation in (4). Thus, we may interpret \u03c4\u03032 as the interactive moment between the links 1 and 2 about the axis Z2. We may also rewrite (4) as \u03c41 = \u03c4\u03031 + \u03c42; \u03c42 = \u03c4\u03032. (9) We have analyzed the dynamic equations of planar manipulators with revolute joints for degrees-of-freedom from 2 to 6 with the help of Maple, a symbolic computational tool by Maplesoft, a division of Waterloo Maple Inc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003238_s12206-021-0829-0-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003238_s12206-021-0829-0-Figure5-1.png",
        "caption": "Fig. 5. Convective heat transfer coefficients of gears and bearing.",
        "texts": [
            " If the residual curve of the parameters in the massconservation, momentum-conservation, and energy-conservation equations tends to be stable, then the simulation has converged. Fig. 4 shows the temperature field contours of the sun gear, planetary gear, ring gear, and planetary gear bearing when the diameter of the planetary gear opening is \u042420 and the number of openings is six (denoted as \u042420\u00d76). The temperature of the inner raceway of the planetary gear bearing is lower than that of the outer raceway due to the influences of the under-ring lubrication and heat generation of the planetary gear. Fig. 5 shows the contours of the convective heat transfer coefficients of key moving components, such as the sun and planetary gears, when the opening parameter of the planetary gear is \u042420\u00d76, and the maximum. The maximum and minimum convective heat transfer coefficients of each key wall can be obtained from the figure, along with the distribution of the convective heat transfer coefficients. On this basis, the CFD method is used to compute the aver- age convective heat transfer coefficient of the key component surface of the gearbox under the following conditions: planetary gear without holes, different diameters of holes, and different numbers of holes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001609_0021998320960531-Figure20-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001609_0021998320960531-Figure20-1.png",
        "caption": "Figure 20. Out of plane compression behaviour of sandwich panels with 60 (left) and 75 (right) corrugated core.",
        "texts": [
            " During the experimental tests the panels were compressed between two stiff plates where the upper FS is restricted to rotate or move along the length of structures so upper FS was constrained in z-direction as shown in Figure 17. The load was applied as displacement on the reference point that was coupled with the upper surface of upper face sheet. Similar approach was also used by Bing et al.34 to investigated the flatwise compression properties of glass fibre reinforced PP corrugated sandwich panels. Rest of the details about the numerical model are discussed in the Bending stiffness of SrPPSBs section. The deflection behaviour of SrPP corrugated structures at different stages under flatwise compression is shown in Figure 20 and Figure 19 along with the mid span deflection d at different points. The first mode buckling is shown by brown arrows, second mode is shown by red arrows while contact of struts with upper FS is shown by white arrows. The load displacement curves for corrugated sandwich structures are shown in Figure 21(a) while more magnified curves with respect to deflection are shown in Figure 21(b). Peak load values are shown in Figure 21(c). The 30 corrugated sandwich structures first show linear behaviour up to d\u00bc 0",
            "05mm the struts comes in contact with the lower surface of upper FS and curve rises up with higher slope as compared to 30 corrugated cores. This can be attributed that struts are shaped into vertical columns that contribute toward load bearing capacity and curve rises up to almost same level. It should be noted here that may be at this stage the struts will undergo plastic hinge failure but our simulation model doesn\u2019t cover the plastic failure of material so this phenomenon cannot be captured during simulations. The deflection behaviour of 60 and 75 corrugated structures is shown in Figure 20. It can be observed from Figure 21(b) that the structures show linear behaviour up to d\u00bc 0.15mm and then nonlinear behaviour of material involves. At d\u00bc 0.41mm the structures undergo first mode of buckling and the load drops. The buckling of strut continues and at d\u00bc 10.5mm, the struts come in contact with the upper FS and load deflection curve moves upward as shown by white arrows in Figure 20. For the structures with 75 corrugation angle the same deflection behaviour is observed. Although the slope of the curve in linear region is higher than those of 60 beams but still the peak load value was lower than 60 beams. As the vertical projected length of struts with the 75 angle is higher as compared to 60 so they exhibited lower peak load because the columns buckle earlier because of higher vertical length of struts which play important role in critical buckling load. In addition to this the struts don\u2019t come in contact with upper FS"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001573_j.ijsolstr.2020.09.004-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001573_j.ijsolstr.2020.09.004-Figure3-1.png",
        "caption": "Fig. 3. Motion of beam axis in Cartesian coordinate frame.",
        "texts": [
            " This implies the necessity of the re-optimization of the adaptive structure under large deformation. This is performed using a machine-learning-based optimization method. The process illustrates how by removing some internal members subject to feasibility constraints, a modified integrated feasible design of the lattice can be obtained which is effective for high-amplitude actuation. In this section, a second-order kinematic and geometrically nonlinear finite element formulation of elastic beams is derived under large displacements and rotations, with the assumption of small strain values. Fig. 3 shows the motion of a beam in a stationary Cartesian coordinate system, where it can experience large displacements and rotations with small strains. The convected orthogonal coordi- aves of variable wave length (Bird et al., 2018). nate system at each point, txi i \u00bc 1;2;3\u00f0 \u00de, is defined along the principle axes of beam at time t, where t e\u03023 denotes the unit tangential vector along the beam axis. In the following, a right subscript i \u00bc 1;2;3 indicates that the variables apply to the convected, rather than the global coordinate system",
            " The main assumptions in this study are as follows: 1. The kinematic conditions do not include warping deformation (i.e., plane sections remain plane); 2. The strains normal to the beam axis are negligible (i.e., the cross section areas remain constant). The kinematics of the beam are studied based on the Updated Lagrangian (UL) incremental approach, which has more numerical effectiveness in comparison with the Total Lagrangian (TL) formulation (Bathe and Bolourchi, 1979). In the UL framework, referring to Fig. 3, the position vector of a generic point P along the beam axis at a known configuration of the beam at time t can be described as tr \u00bc xt3e\u03023 \u00f01\u00de Then, the position vector of a point in the corresponding cross section is expressed as tR \u00bc tr \u00fe xt1e\u03021 \u00fe xt2e\u03022 \u00f02\u00de During motion of the beam to an unknown configuration at time t \u00fe Dt, making use of assumptions 1 and 2, this same point deforms to t\u00feDtR \u00bc t\u00feDtr \u00fe x1t\u00feDt e\u03021 \u00fe x2t\u00feDt e\u03022 \u00f03\u00de where t\u00feDtr \u00bc tr \u00feW , in which W x3\u00f0 \u00de \u00bc w1 w2 w3f gT expresses the incremental displacement vector of the beam axis and is measured in the convected coordinate system at time t. If the incremental finite rotation components around the principle axes of beam t e\u0302i at the generic point P in Fig. 3 are denoted with H x3\u00f0 \u00de \u00bc h1 h2 h3f gT , the updated director vectors after some geometric manipulations can be expressed as t\u00feDt e\u0302i \u00bc H t e\u0302i \u00fe 1 2 H H t e\u0302i \u00f04\u00de Further details concerning this relation can be found in Argyris (1982). Using Eq. (4), the new transverse director vectors are obtained as t\u00feDt e\u03021 \u00bc 1 1 2 h22 \u00fe h23 h3 \u00fe 1 2 h1h2 h2 \u00fe 1 2 h1h3 T \u00f04a\u00de t\u00feDt e\u03022 \u00bc h3 \u00fe 1 2 h1h2 1 1 2 h21 \u00fe h33 h1 \u00fe 1 2 h2h3 T \u00f04b\u00de Substituting 4 in Eq. (3) and subtracting Eq. (1) from the result, yields the following second-order approximations for increments in the displacement field with finite rotations from time t to time t \u00fe Dt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure8-1.png",
        "caption": "Fig. 8. Topology optimization outcome, for \u03c1>0.85; top view (top) and 3D view (bottom)",
        "texts": [
            " (1) and its boundary conditions are fulfilled along with an upper limit on total material used, i.e. max 0<\u03c1\u22641 { 1 2 \u222b \u2126 D(\u03c1(r))TTT d\u2126 } s.t. : \u2207 \u00b7 S = 0, (5)\u222b \u2126 \u03c1(r)d\u2126 \u2264 f \u00b7 V, where V is the volume of the domain \u2126 and 0 < f \u2264 1 is the maximum volume fraction allowed. To avoid the checkerboard problem [26] we employ a volume-preserving Helmholtz filter with a constant radius r0, and to avoid intermediate values of \u03c1, we employ a tanhtype projection function [27] in two iterations with slopes \u03b2 = 1 followed by \u03b2 = 8, respectively. Figure 8 illustrates the latter case in the form of an .stl file. In order to validate the structure of the new geometry, the stress was investigated. The maximum von Mises stress experienced by the optimized part for f = 0.5 is approximately 36MPa, slightly below the 3D printing PLA ultimate tensile strength [28]. Further reduction of f would therefore increase the maximum stress to an unacceptable level. The topology optimization has proven successful in producing a geometry which offers the same stiffness at a fraction of the original material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000220_ecce.2019.8913300-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000220_ecce.2019.8913300-Figure7-1.png",
        "caption": "Fig. 7. Configuration of Compared Machine.",
        "texts": [
            " The prototype is under manufacturing, and will be tested soon in the future. (a) (b) (c) IV. PERFORMANCES COMPARISON BETWEEN MOTORS WITH HALBACH PM ARRAY AND NORMAL SECONDARY To investigate the influence of Halbach magnet array to the proposed machine\u2019s performance, a structure with a regular bipolar secondary is designed for comparison. Both dimension parameter of primary mover module and current density are set to the same to validate the comparison. Structure of the compared motor is shown in Fig. 7. 3-phase modular TC-LPMVMs with regular secondary and Halbach PM array secondary are both analyzed through 3-D finite element analysis. There is a great improvement on the average thrust force. (a) -30 -20 -10 0 10 20 30 0 60 120 180 240 300 360 E M F( V ) elec. degree -100 0 100 200 300 400 500 600 0 60 120 180 240 300 360 T hr us t f or ce ( N ) elec. degree A B C D E Total -200 0 200 400 600 800 1000 1200 0 60 120 180 240 300 360 T hr us t f or ce ( N ) elec. degree A B C D E Total 200 250 300 350 400 0 60 120 180 240 300 360 T hr us t f or ce ( N ) elec"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001242_tec.2020.3007715-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001242_tec.2020.3007715-Figure10-1.png",
        "caption": "Fig. 10: Prototyped motor (a) stator and (b) rotor dimensions (mm).",
        "texts": [],
        "surrounding_texts": [
            "For the measured load performance, the prototyped test motor is synchronized with the 400 V grid supply and loaded by varying the dynamometer power in the test bench of Fig. 12. For these tests, the field current was set constant I f = 23 A and the winding temperature monitored at 75 \u25e6C. The measured load performance results as a function of input power are shown in Fig. 15. Authorized licensed use limited to: University of Exeter. Downloaded on July 16,2020 at 00:15:25 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. To predict the load performance, the iterative process of Fig. 6 is used in which the grid load angle is calculated using the mapping technique given in Section V. The performance curves of power factor, efficiency, power input, power output and current are predicted by (12), (14), (15), (20) and (27) respectively. The predicted load performance curves as a function of input power are shown in Fig. 15. This shows good agreement between predicted and measured load performance. The slight difference in the power factor and current in Fig. 15(b) at low input powers can be explained by the FEM used BH curve. This aspect is further considered in Sub-section VIII-A."
        ]
    },
    {
        "image_filename": "designv11_63_0001689_1754337120961609-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001689_1754337120961609-Figure8-1.png",
        "caption": "Figure 8. The three seam orientations used in this study. The \u2018\u2018Top Loop\u2019\u2019 orientation has a seam that is always near the top spin pole. The \u2018\u2018Bottom Loop\u2019\u2019 has a seam that is always near the bottom spin pole. The 2-seam has no seam near the hemisphere plane.",
        "texts": [
            " The key to Seam Shifted Wake pitches is to allow the deflection to remain on the same side of the ball for most of a rotation. While there are likely many orientations that can produce a Seam Shifted Wake on a spinning pitch, we must work within some constraints. First, while the cannon and flex tip can project a ball with a wide range of spin rates and axes, it cannot impart a spin component in the Y direction, commonly called \u2018\u2018gyro\u2019\u2019 spin. A simple candidate for a Seam Shifted Wake pitch with no gyro spin component is an orientation that we term \u2018\u2018Looper\u2019\u2019 and is shown in Figure 8. The name comes from the seam looping around the pole as the ball spins. For most of the rotation, that seam remains within the angles that cause early separation as shown in Figure 7. Three orientations were tested as shown Figure 8. Two of the orientations (called Top Loop and Bottom Loop) are designed to generate seam effects. The third, which is a standard 2-seam orientation, should not experience such effects and will be used as the baseline. While the spin axis was the same for all pitches, different orientations (relative to the spin axis) and seam heights were used. The tests consisted of 12 pitches in the Top Loop and Bottom Loop orientation for two baseballs with different seam heights. Additionally, four pitches in the 2-seam orientation using the smaller-seamed ball form a baseline"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000162_012003-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000162_012003-Figure5-1.png",
        "caption": "Figure 5. Stress distribution of the oversized model. Figure 6. Optimization result.",
        "texts": [
            " From the two analyses, the first was the most critical since it had higher stress and displacement values, see Figure 4(a) and Figure 4(b). The equivalent von Mises stress distribution resulted from the first analysis had a maximum value of approximately 22 MPa, and the maximum total displacement was approximately 5E-5 m. As previously mentioned, in order to carry out the process of shape optimization, an oversized control arm was designed to later remove the zones that presented low stresses. Figure 5 shows the equivalent von Mises stress of this oversized part, the maximum value was near 3 MPa. To achieve the optimal shape, it was necessary to perform 29 iterations. The result obtained from the optimization process is shown in Figure 6. 5th IMRMPT Journal of Physics: Conference Series 1386 (2019) 012003 IOP Publishing doi:10.1088/1742-6596/1386/1/012003 The geometry resulting from the topology optimization was post-processed to obtain a non-faceted reconstructed geometry, which facilitates the design and manufacture of the final prototype"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure7-1.png",
        "caption": "Fig. 7 Boundary conditions of the output bearing seat and the inner bracket 1",
        "texts": [
            " The transmission force can be calculated as follows: Consequently, the force transmission rate of the nth path is: The contribution of the vibration path of the double-layer box is mainly affected by the magnitude of the stiffness and damping between the parts that are in contact with each other; thus, the stiffness and damping need to be determined. The output bearing seat is in direct contact with the inner bracket 1 in the radial direction. When the output bearing seat is subjected to vibration, the vibration is transmitted to the inner bracket 1 through the contact surface. Figure\u00a07 presents the boundary conditions of the output bearing seat and the inner bracket 1. The contact type between the output bearing seat and the inner bracket 1 surfaces was set to bonded contact. The outer surface of the inner bracket 1 was defined as a fixed support, and a load F ranging from 100 to 500\u00a0N was applied to the inner surface of the output bearing seat along the Y direction. Figure\u00a08 shows the radial static displacement of the output bearing seat and the inner bracket 1 under load F, based on which, the deformation of the contact surface can be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure17-1.png",
        "caption": "Fig. 17. CPM machine with (a) existing step-staggered rotor containing none magnetic barriers (CPM-II-A), (b) proposed unequal length step-staggered rotor containing magnetic barriers (CPM-II-B), and (c) proposed multistep step-staggered rotor containing magnetic barriers (CPM-II-C).",
        "texts": [
            " Thus, the magnetic field is divided into the radial and axial magnetic fields. The radial flux is coupled with the windings to generate back-EMFs. However, the axial flux may increase the unipolar leakage flux in the end region, as shown in Fig. 16 (b2) and (b3), which would affect the bearings and position sensors. Therefore, the axial leakage flux in the end region is inevitable and cannot be negligible, and the magnetization risk of end shaft in the CP PM machine is much higher than that in the PM machine with N-S pole rotor. Fig. 17 (a) shows the existing staggered machine [28], [29]. It is composed of two staggered rotors, the axial length of the staggered segments is equal and there are no magnetic barriers between the staggered segments. Therefore, there is significant amount of interpole axial leakage flux not coupling with the armature windings. To reduce the interpole axial leakage flux, two unequal and multistep staggered rotors are proposed, as shown in Fig. 17 (b) and (c), respectively. Magnetic barrier rings are introduced between staggered stators and between staggered rotors. In addition, a magnetic barrier cylinder is embedded between the inner surface of the rotor and the outer surface of the shaft (made of stainless steel). The aluminum is selected as low permeance material for this study. The staggered angle between the two adjacent rotors is \u03c0/Pr mechanical degrees, which is 180 electrical degrees (one pole pitch). This means that the magnetization direction of PMs is opposite in the adjacent rotors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure14-1.png",
        "caption": "Fig. 14 Push-pulling the motorcycle triple clamp model and the modeling results",
        "texts": [
            " Push-pulling the triple clamp model in Fig.\u00a013 involved 10 critical points in total, and some of them occurred concurrently. As a result, the GTI configurations are very complex, and the associated detection task is challenging. Siemens NX was only able to successfully cross the first two critical points, and SpaceClaim was not even able to cross a single critical point (yielding a weird resulting geometry), while the proposed method can correctly detect all the critical points and resolve the generated GTI. In case study 5 (Fig.\u00a014), instead of linearly adding more critical points to the push\u2013pull move, the comprehensiveness was attained by push-puling a same model under various situations: (1) push\u2013pull the blue faces and stop in between the first and second critical pointsz; (2) push\u2013pull the blue faces, and stop in between, then continue the push\u2013pull until the end; and (3) push\u2013pull the blue faces until the end. The modeling result for the first situation is shown in the middle of the upper row in Fig.\u00a014, that for the second situation is shown in the upper-right, and that for the third situation is the same as the second one. Siemens NX failed to update the model for all of the three situations, while the proposed method can successfully update the model for all of them. The comparisons in case studies 1\u20135 are sufficient to show that the proposed method outperforms the state of the art in terms of robustness. Nevertheless, only translational push-pulls were used in these case studies, and the push\u2013pull ranges were small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001088_b978-0-12-819972-5.00006-9-Figure6.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001088_b978-0-12-819972-5.00006-9-Figure6.3-1.png",
        "caption": "Figure 6.3 Components of drones. (From Ref. [17].)",
        "texts": [
            " If these variables are not properly selected, the following results may bring about negative effects [15]. These simulators are used to create mathematical models that may be complex and extremely similar to the main system, thereby reducing the experimental cost and time. 3 Where it is applicable Drones are built with several sensors and hardware for different purposes and for this particular reason, they are considered as an integral part of IoT. They both comprise of AI, sensing, data/information processing, and communication networks [16] (Fig. 6.3). These sensors send and receive data through AI techniques embedded in their systems, if the systems are actually built with AI algorithms (United States Patent No. US 2019/0176987 A1, 2019) and perform the required operation afterwards. Usually, we simulate some of the following components of a complete drone (Figs. 6.4 and 6.5): \u2022 Vision and control Drones perform vision through optical sensing or cameras. This way, they perceive and understand their environment and detect objects for proper navigation [19]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002138_s11668-020-01106-2-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002138_s11668-020-01106-2-Figure2-1.png",
        "caption": "Fig. 2 Numerical model of composite cylindrical roller and its meshing",
        "texts": [
            " Initial and Boundary Conditions Since the composite cylindrical roller bearing is an axisymmetric structure, and the loading condition and the structure of the cylindrical roller are also symmetric, the numerical model of the composite cylindrical roller can be simplified by taking 1/2 bearing roller, 1/4 bearing outer ring and 1/16 bearing inner ring to establish a numerical calculation model. The outer surface of the outer ring of the bearing is fixed, the symmetry plane of the bearing is symmetrically restricted, and the section is freely restricted. Using hexahedron to mesh the numerical model, the numerical model of the composite cylindrical roller is shown in Fig. 2. Comparison with Hollow Cylinder Roller Applying a light load of 15 kN and a heavy load of 25 kN to the hollow cylindrical roller and composite cylindrical roller, respectively, analyzes and compares the force of the two rollers under the two load conditions. The analysis and comparison results are shown in Table 1. Under the load condition of 15 kN, the maximum equivalent stress reduction rate of the composite cylindrical roller bearing is 10.8%, the maximum contact stress reduction rate is 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001968_me49197.2020.9286694-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001968_me49197.2020.9286694-Figure6-1.png",
        "caption": "Fig. 6. Coupled model",
        "texts": [
            " Downloaded on June 18,2021 at 23:46:29 UTC from IEEE Xplore. Restrictions apply. IV. COUPLED MODEL WITH 1-DOF MANIPULATOR The analysis presented in the previous section has shown that the dynamics of the roll, pitch, and yaw axes are essentially decoupled. Therefore, the development of a planar model in the vertical plane ( ) is meaningful. This model is developed to study the main features of the coupling between UAV dynamics and manipulator dynamics. A robot arm having 1 DOF is considered. Therefore, the model, which is represented in Fig. 6, has 4 DOFs, which are associated to the coordinates of the CM of the UAV ( and ), roll angle \u03d5, and rotation \u03b8 of the robot arm with respect to the UAV. The equations of motion of the coupled system are derived with Lagrange\u2019s method. The kinetic ( ) and potential energy of the system are given by the following equations: = 12 ( + ) + 12 \u03d5 + 12 + 12 \u03d5 + \u03b8 = + , (5) in which (= ) is the moment of inertia of the UAV about its CM, and are the mass and moment of inertia of the manipulator about its CM, and , are the velocity and vertical position of the robot CM, and is the gravitational acceleration. If the position and velocity of the robot CM are expressed as a function of the selected coordinates, the following equations are obtained (see Fig. 6 for the definition of and ): = 12 + 12 + 12 \u03d5+ 12 + + \u03d5 + 2 (\u03d5)\u03d5+ 2 (\u03d5)\u03d5 + 4 \u03d5 + 4 \u03b8 + 2 + 4 \u03d5\u03b8+ 2 2 (\u03d5 +\u03b8)\u03d5 + 2 2 (\u03d5 +\u03b8)\u03d5+ 2 2 (\u03d5 +\u03b8)\u03b8+ 2 2 (\u03d5 +\u03b8)\u03b8+ 2 2 (\u03d5)\u03d5 + 2 2 (\u03d5)\u03d5\u03b8 + 12 \u03d5+ 12 \u03b8 + 12 \u03d5\u03b8 = + \u2212 (\u03d5) \u2212 2 (\u03d5 +\u03b8) (6) Lagrange\u2019s equations are: dd \u2202\u2202 \u2212 \u2202\u2202 = (7) In (7), = \u2212 is the Lagrange\u2019s function, are the generalized coordinates ( , , \u03d5 , and \u03b8 ), and the generalized forces. The introduction of (6) in (7) leads to the dynamics equations of the coupled system: + + (\u03d5) + 2 (\u03d5 + \u03b8) \u03d5 + 2 (\u03d5 + \u03b8)\u03b8 + \u22122 (\u03d5 + \u03b8)\u03b8 + (\u03d5) + 2 (\u03d5 + \u03b8) \u03d5 \u2212 22 (\u03d5 + \u03b8)\u03d5\u03b8 = \u2212 (\u03d5) (8) + + (\u03d5) + 2 (\u03d5 + \u03b8) \u03d5 + 2 (\u03d5 + \u03b8)\u03b8 + +2 (\u03d5 + \u03b8)\u03b8 + (\u03d5) + 2 (\u03d5 + \u03b8) \u03d5 \u2212 22 (\u03d5 + \u03b8)\u03d5\u03b8 + += (\u03d5) (9) (\u03d5) + 2 (\u03d5 + \u03b8) + (\u03d5) + 2 (\u03d5 + \u03b8) + + + + 4 + 2 2 \u03d5+ 2 (\u03b8)\u03b8\u2212 2 (\u03b8) \u03d5 + 2 2 (\u03b8)\u03d5\u03b8+ 2 (\u03b8)\u03d5+ (\u03b8) + 2 (\u03d5 + \u03b8) = (10) 2 (\u03d5 + \u03b8) + 2 (\u03d5 + \u03b8) + + 4 + 2 (\u03b8) \u03d5 + + 4 \u03b8 + 2 (\u03b8)\u03d5+ 2 (\u03d5 + \u03b8) = (11) It is worth noticing that these equations hold true even for large rotations of the UAV and of the manipulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001881_icca51439.2020.9264580-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001881_icca51439.2020.9264580-Figure1-1.png",
        "caption": "Figure 1. Quadcopter\u2019s model",
        "texts": [
            " performance of UAVs is significantly enhanced by using the previously mentioned methods, the chattering effect is still existent due to a discontinuous term of the switching control law. The motivation of this study is to design a superior controller with high quality robustness, rapid adaptation without chattering effect on a quadcopter UAV. II. QUADROTOR MODELING The mathematic model of a quadrotor is introduced in many existent approaches [15-18]. The basic frame consists of an Earth frame, {E}, and body frame, {B}, as shown in Figure 1. Let , ,  , ( 2 2\u2212   , 2 2\u2212   , )\u2212   for all time, represent the Euler angle (roll, pitch and yaw), respectively; , ,x y z  are the position in the {E}. iF (i=1,2,3,4) denotes the thrust generated by motors i; i is the speed of the rotor i; , , are external disturbances influencing on roll, pitch, yaw dynamics of the vehicle, respectively. In this study, the external disturbance is considered as a harmonic disturbance with known frequency but unknown amplitude. The dynamics model of the vehicle including harmonic external disturbance ( , , ) can be described as follows equations [15]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000729_humanoids43949.2019.9035004-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000729_humanoids43949.2019.9035004-Figure16-1.png",
        "caption": "Fig. 16. Schematic of the hardware used to interface with the RPAs in the helicon plasma source. The housing of the conventional RPA is identical to that of the hybrid RPA and hence they used the same mount.",
        "texts": [
            " 15): (i) a low-power, capacitively coupled mode (E mode) characterized by a hollow plasma, with larger number density in a halo roughly the size of the antenna diameter; (ii) a mid-power, inductively coupled mode (H mode) with roughly uniform distribution within the antenna diameter and a drop in plasma potential compared to the E mode [39]; and (iii) a high-power, helicon mode (W mode) distinguished by a conical shape [40]. A large jump in density occurs between the capacitive and inductive modes, and between the inductive and helicon modes. Each jump in plasma density is associated with an increased brightness of the discharge [39]. The experimental procedure involves first installing the RPA under study in an arm with electrical feedthroughs that connect the sensor to instrumentation installed outside the chamber (Fig. 16). The chamber is then pumped to a base pressure of about 2\u00d710\u22125 Torr before helium is flowed into the system until a pressure of approximately 5 Pa (about 4\u00d710\u22122 Torr) is reached. The confining magnetic field was set to 500 Gauss. Data were collected over a wide range of antenna RF power (at a frequency of 13.56 MHz) using the various RPAs as well as a double Langmuir probe previously installed in the chamber. In the experiments, the bias voltage of the electron-repelling grids was set at \u221250 V"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001271_s12541-020-00378-w-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001271_s12541-020-00378-w-Figure2-1.png",
        "caption": "Fig. 2 LIPM",
        "texts": [
            " The formulated problem becomes a finite-time horizon discretetime linear quadratic regulator (LQR) problem for a time-varying system because it takes into account the vertical acceleration of the CoM. This problem can be solved in finite time by the backward process method, which is described in Algorithm\u00a01 in Sect.\u00a03. To evaluate the proposed algorithm, experiments comparing it with an existing method as well as evaluations of its calculation time and its application in a biped robot simulation were conducted. The results are reported in Sects.\u00a03 and 4.3. Section\u00a05 concludes the paper. A biped humanoid robot can be represented approximately as an LIPM [9]. As illustrated in Fig.\u00a02, an LIPM represents the relationship between the position of the CoM, denoted by c = [ cx;cy;cz ] \u2208 \u211d 3 , and the position of the ZMP, denoted by p = [ px;py;0 ] \u2208 \u211d 3. Let the angular momentum at c with respect to the x-and y-axes be Lx and Ly,respectively. Provided that there are no external forces, the position of the ZMP is given by where m is the total mass of the robot and g is gravitational acceleration. Supposing that the derivative of the angular momentum is sufficiently small to be neglected, the above ZMP equation reduces to and this can be written as the following linear time-varying system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002411_iros45743.2020.9340935-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002411_iros45743.2020.9340935-Figure7-1.png",
        "caption": "Fig. 7: Gait design of the quadruped robot. (a) Definition of the rotating angle \u03b1 in inchworm robot. (b) Notation of the quadruped robot. (c) Control signals of the eight inchworm robots. The combined motions of the inchworm robots make the quadruped robot perform trot gait walking.",
        "texts": [
            " We used a master computer to communicate with all eight inchworms and control the gait of the quadruped robot. The architecture of the gait control consists of four levels (Fig. 6). The first level is gait selection, where the gait of the quadruped robot is selected. In this work, we adopt the trot gait of the quadruped robot as an example. Based on the selected trot gait, the second level (motion generator) produces eight individual rotation cycles for each joint of the quadruped robot. Rotation cycles of each joint for trot gait are predefined in the master computer (Fig. 7(c)). Then the master computer sends all the rotation cycles to the receivers of the inchworm robots (Fig. 7(b)). The third level takes place on the onboard controller of the inchworm robots. After receiving the rotation signals, the micro-controllers of the individual inchworm robots serve as CPGs to generate periodic PWM signals based on the joint rotation signals. The transfer relationship is predefined in the onboard controller of the modules. The servos execute the PWM signals to ensure that the whole quadruped robot performs a periodic trot gait. In this part, we characterize the metrics of individual modules including their velocity and turning radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001197_s0263574720000533-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001197_s0263574720000533-Figure13-1.png",
        "caption": "Fig. 13. The head-raising motion by the predefined spiral curve method.",
        "texts": [
            " University of Exeter, on 03 Jul 2020 at 02:30:12, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. \u23a7\u23aa\u23a8 \u23aa\u23a9 t\u03021 = \u2211n i=1 li t\u03022 = t\u03021 + \u03c6base t\u03023 = t\u03021 + 2nc\u03c0 (46) where a, b, and c are adjustable parameters. \u03c6base is the phase value of the second interval, and nc is the cycle number of the spiral curve. And Eq. (43) represents the straight line segment, and Eq. (44) describes the planar spiral curve segment. And Eq. (45) represents the spatial spiral curve segment. The predefined spiral curve is shown in Fig. 13. Then the snake robot moves from state 1 to state 2, and the pose of the head forms a spatial trajectory. The parameters of the predefined curve are as follows: a = 14.7, b = 1, c = 48.5, t1 = 1552, t2 = 1555.1, t3 = 1567.7, \u03c60 = 1.6022, \u03c6base = \u03c0 , and nc = 2.5. The motion process of the head-raising for the snake robot is shown in Fig. 13. The starting state in which the snake robot raises its head is the state 1 as shown in Fig. 13. And the end state in which the snake robot raises its head is the state 2. The workspace of the method is the spatial trajectory formed by the pose of the head. 3.1.2. The workspace of the method based on the B\u00e9zier curve. In order to get the workspace of the head for the snake robot, we are inspired by the method in ref. [33]. Firstly, n0 grid points are evenly distributed in the hemisphere space whose center point is P0 and radius r0 equals the length of the upper part of robot, \u221117 i=9 li. P0 is the end point of the base part of the snake robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000295_icems.2019.8921961-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000295_icems.2019.8921961-Figure2-1.png",
        "caption": "Fig. 2. Temperature Cloud of Hub Motor under Natural Cooling",
        "texts": [
            " With the help of fluent software, the temperature field of the motor is simulated. The fluid field is set to wall and the ambient temperature is set to 25 \u00b0C. The surface of the parts that need to be given heat dissipation coefficient is presented separately. According to Table 4 and Table 5, the heat dissipation coefficient and thermal conductivity of each part of the motor are given respectively. The steady-state temperature field cloud of each part of the hub motor under natural cooling is shown in Figure 2. The maximum temperature of each part of the motor under natural cooling is shown in Table 6. According to the temperature field cloud of each part of the motor under natural cooling in Fig. 2, it can be seen that the winding of the motor is the most heat-intensive part of the whole generator. From Table 6, it is observed that the maximum temperature of the winding reaches 184.19\u00b0C. The highest temperature is in the middle of the winding in the slot, and the lowest temperature is at the end of the winding. According to the thermal conductivity of the material in Table 4, the thermal conductivity of winding insulation is low and the insulation effect is good. The heat generated by copper winding in the slot is difficult to transfer through insulation, so the temperature of the winding in the slot is higher"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure21-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure21-1.png",
        "caption": "Fig. 21 The schematic diagram",
        "texts": [
            " In this mechanism, spring is added to kinematic joint G, so that the slider 5 and the component 6 are combined into one component under the constraint of spring, and the nipper is in an open state at this time. The nipper closed configuration is shown in Fig.\u00a020b. In this mechanism, a geometric constraint is added at kinematic joint C, and the upper nipper 4 which combined with the lower nipper 2 to form one component is gradually moved from position b to position a. Based on the principle of augmented Assur groups, the metamorphic nipper swing mechanism is divided into a fixed-axis rotating active part, an Assur group RPR and an augmented Assur group RP\u2013RR\u2013RR\u2013R, as shown in Fig.\u00a021. According to geometric and inertia properties of the metamorphic nipper swing mechanism in Table\u00a0 3, a three-dimensional model is established in SolidWorks, as shown in Fig.\u00a022. The initial distance between slider 5 and H is 172.37\u00a0mm, and the initial position of the component 9 is 3 \u22155\u00a0rad. Assuming that the active part 9 rotates at a constant speed of 6 r/min (motion period is 10\u00a0s), the dynamic simulation is carried out in SolidWorks virtual prototype environment, and the relationship between the driving torque and time of the metamorphic nipper swing mechanism is obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002521_j.euromechsol.2021.104267-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002521_j.euromechsol.2021.104267-Figure18-1.png",
        "caption": "Fig. 18. Equivalent designs assuming \u03940\u03d11 = \u2212 \u03940\u03d12.",
        "texts": [
            " This physical effect happens because the spring remains on an unstable static equilibrium state at rest and, if relative motion occurs, a resulting force provides torque in favor of it. For \u03bbi = 0.5 (Fig. 16a), the maximum amplitudes are slightly different from the condition with tensioned springs (Fig. 14a), as the peak value is reduced by half and no jumps were detected. On the other hand, for \u03bbi = 0.7 (Fig. 16b), there is a reduction of the response. According to Fig. 17a, jumps are localized within a narrow frequency range for \u03b6 = 0.01, and the system jumps-down no matter the sweeping direction adopted. The saddle-node bifurcation is verified in Fig. 17b. Although Fig. 18a describes the case where both springs are pinned at the same point on the outer track, the resulting torque is equivalent to the condition on Fig. 18b. For similar properties, arrangements with two elements produce the same torque curve if \u03940\u03d11 = \u2212 \u03940\u03d12. The designs in Fig. 18c and d are produced if one of the springs is shifted by 180\u2218. H.H. Miyasato et al. European Journal of Mechanics / A Solids 89 (2021) 104267 If the angles are adopted as \u03940\u03d11 = 5\u2218 and \u03940\u03d12 = \u2212 5\u2218 (\u03c11 = \u03c12 = 1.0), the magnitudes of the non-dimensional torque in Fig. 19a nearly double the figures obtained using a single element (Fig. 4a). If necessary, the torque obtained without pre-existing static forces can be adjusted by a reduction on \u03c1i. Besides, Fig. 19b shows that this is another possibility of introducing a non-zero stiffness stage for \u0394\u03b8 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure3-1.png",
        "caption": "Fig. 3. (a) Analytical model forces and deformations; (b) Degrees of freedom of the MATRIX27 element.",
        "texts": [
            " Finally, the results of these efficient FE models will be shown, and compared with a FE reference analysis for validation purposes. For the sake of clarity and completeness of the work, the analytical formulation in [4] will be summarized here. As mentioned before, and illustrated in the upper part of Fig. 2, the formulation reproduces the structural behaviour of the wire-roller-wire subset with an equation system. In particular, a set of 6 equations is needed, which relate the parameters shown in the representation of the analytical model in Fig. 3a. The basis of this formulation lies on assuming that the roller-wire contact remains in stick status and the wire-ring contacts in slip status. This way, the wire-roller-wire subset behaves as a rigid solid. N2 = k2\u22c5\u22062 (2) (N2 \u00b1 \u03bc\u22c5N1)\u22c5cos(\u03b1)+ (N1 \u2213 \u03bc\u22c5N2)\u22c5sin(\u03b1) = k3\u22c5\u22063 (3) DCW/2\u22c5cos(\u03b10) = \u2206R/2 \u2212 \u22062 +(DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) (4) DCW/2\u22c5sin(\u03b10) = \u2206A/2 \u2212 \u22061 +(DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) (5) N2\u22c5((DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) ) \u00b1 \u03bc\u22c5N1\u22c5(((\u03bb/2 \u2212 \u22061) + (DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) ) \u00b1 \u03bc\u22c5N2\u22c5((\u03bb/2 \u2212 \u22062) + (DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) ) \u2212 N1\u22c5((DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) ) = 0 (6) According to Fig. 3a, \u0394A and \u0394R are the input axial and radial displacements, respectively (caused by external loads). These displacements give rise to normal contact forces N1 and N2 and their respective deformations \u03941 and \u03942 in the wire-ring contact zones. Similarly, the roller-wire contact also gets deformed (\u03943) and its normal contact angle varies from \u03b10 to \u03b1. Finally, Dcw is the distance between wire centres, \u03bb is the wire diameter, and \u03bc is the coefficient of friction. Contact forces and displacements are linked via the contact stiffness constants (k1, k2, k3), whose calculation will be introduced later",
            " 1\u20136 depend on which input displacement (\u0394A or \u0394R) is the prevailing one [4]; the upper signs correspond to prevailing \u0394A and the lower signs to prevailing \u0394R. This behaviour is related to the wire twisting phenomenon [9], because the wire twists in one direction or in the other, depending on the prevailing relative displacement between rings, generating friction forces in opposite directions for each case. To sum up, the analytical formulation provides the contact angle, forces and deformations, caused by axial and radial external displacements, as shown in Fig. 4. For type B rollers (see Fig. 1), the analytical model of Fig. 3a must be vertically flipped. Equivalent Eqs. 1-6 are obtained by applying the same reasoning, where only the sign of N1 changes. In order to simplify the explanations, mathematical developments are based on the type A roller case from now on. In order to implement Eqs. 1-6 equation system in Ansys\u00ae, MATRIX27 element was used, which allows to introduce a linear userdefined matrix that links forces and displacements between two points with six degrees of freedom (\u03b4x1, \u03b4y1, \u03b4z1, \u03b4x2, \u03b4y2, \u03b4z2), as illustrated in Fig. 3b. However, this is not straightforward, because Eqs. 1-6 form a Wire Outer Ring Inner Ring Roller A Roller B Fig. 1. Crossed roller wire bearing cross section [4]. I. Mart\u0301In et al. Tribology International 161 (2021) 107098 non-linear equation system which is not related with the MATRIX27 degrees of freedom. Consequently, two additional steps are needed, one to transform the non-linear equation system defined in Eqs. 1-6 into a linear one, and another one to relate them with the degrees of freedom of the two reference points",
            "1 = \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 0 b c 0 j \u2212 1 d e k 0 f g 0 i 0 0 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 = j\u22c5( \u2212 c\u22c5g\u22c5i) + k\u22c5(i\u22c5c\u22c5e) = Det.1j\u22c5j + Det.1k\u22c5k Det.2 = \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 a 0 c 0 0 j d e \u2212 1 k f g h 0 0 0 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 = j\u22c5(c\u22c5g\u22c5h)+ k\u22c5( \u2212 h\u22c5c\u22c5e) = Det.2j\u22c5j+Det.2k\u22c5k Det.3 = \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 a b 0 0 0 \u2212 1 j e \u2212 1 0 k g h i 0 0 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 = j\u22c5(a\u22c5g\u22c5i \u2212 b\u22c5g\u22c5h)+ k\u22c5( \u2212 a\u22c5e\u22c5i+ b\u22c5e\u22c5h) = Det.3j\u22c5j+Det.3k\u22c5k Det.4 = \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 a b c 0 0 \u2212 1 d j \u2212 1 0 f k h i 0 0 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 \u20d2 = j\u22c5(h\u22c5b\u22c5f \u2212 i\u22c5a\u22c5f \u2212 i\u22c5c)+ k\u22c5( \u2212 h\u22c5b\u22c5d \u2212 h\u22c5c+ i\u22c5a\u22c5d) = Det.4j\u22c5j+Det.4k\u22c5k According to Fig. 3, the following change of variables can be carried out: j = \u2212 \u2206R 2 = \u2212 1 2 \u22c5(\u03b4x2 \u2212 \u03b4x1) (18) k = \u2212 \u2206A 2 = \u2212 1 2 \u22c5 ( \u03b4y1 \u2212 \u03b4y2 ) (19) Where \u03b4x1,\u03b4y1,\u03b4x2 and \u03b4y2 are the degrees of freedom of the MATRIX27 element. Thus, Eqs. 1\u20132 can be rewritten as: Ni 1 = k1\u22c5\u22061 = k1\u22c5 Det.1 Det. = k1\u22c5 Det.1j\u22c5 ( \u2212 1 2 \u22c5(\u03b4x2 \u2212 \u03b4x1) ) + Det.1k\u22c5 ( \u2212 1 2 \u22c5 ( \u03b4y1 \u2212 \u03b4y2 )) Det. (20) Ni 2 = k2\u22c5\u22062 = k2\u22c5 Det.2 Det. = k2\u22c5 Det.2j\u22c5 ( \u2212 1 2 \u22c5(\u03b4x2 \u2212 \u03b4x1) ) + Det.2k\u22c5 ( \u2212 1 2 \u22c5 ( \u03b4y1 \u2212 \u03b4y2 )) Det. (21) Again, according to Fig. 3: Fx1 = \u2212 N2 \u2213 \u03bc\u22c5N1 (22) Fy1 = N1 \u2213 \u03bc\u22c5N2 (23) Fx2 = N2 \u00b1 \u03bc\u22c5N1 (24) Fy2 = \u2212 N1 \u00b1 \u03bc\u22c5N2 (25) Replacing Eqs. 20\u201321 in Eqs. 22\u201325: Fx1 = \u2212 k2\u22c5 Det.2j\u22c5 ( \u2212 1 2 \u22c5(\u03b4x2 \u2212 \u03b4x1) ) + Det.2k\u22c5 ( \u2212 1 2 \u22c5 ( \u03b4y1 \u2212 \u03b4y2 )) Det. \u2213 \u03bc\u22c5k1\u22c5 Det.1j\u22c5 ( \u2212 1 2 \u22c5(\u03b4x2 \u2212 \u03b4x1) ) + Det.1k\u22c5 ( \u2212 1 2 \u22c5 ( \u03b4y1 \u2212 \u03b4y2 )) Det. (26) \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u00b1\u03bc\u22c5k1\u22c5cos(\u03b10) +k1\u22c5sin(\u03b10) k2\u22c5cos(\u03b10) \u2213\u03bc\u22c5k2\u22c5sin(\u03b10) \u2212 k3 0 0 \u2212 1 \u2212 cos(\u03b10) \u2212 (DCW/2)\u22c5sin(\u03b10) \u2212 1 0 \u2212 sin(\u03b10) (DCW/2)\u22c5cos(\u03b10) \u00b1\u03bc\u22c5k1\u22c5\u03bb/2 \u00b1\u03bc\u22c5k1\u22c5DCW/2\u22c5sin(\u03b10) \u2212 k1\u22c5DCW/2\u22c5cos(\u03b10) k2\u22c5 DCW 2 \u22c5sin(\u03b10) \u00b1\u03bc\u22c5k2\u22c5\u03bb/2 \u00b1\u03bc\u22c5k2\u22c5DCW/2\u22c5cos(\u03b10) 0 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u22c5 \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u22061 \u22062 \u22063 \u2206\u03b1 \u23ab \u23aa \u23ac \u23aa \u23ad = \u23a7 \u23aa \u23a8 \u23aa \u23a9 0 \u2212 \u2206R/2 \u2212 \u2206A/2 0 \u23ab \u23aa \u23ac \u23aa \u23ad (16) I",
            " 10a shows the wire-ring contact status when solid wires are used in the simulations, being coloured in orange the regions in contact. Fig. 10b and 10c show, coloured in grey, the surfaces of the rings to which the MATRIX27 and COMBIN39 are linked, respectively. These surfaces are rigidized though shell elements, so no local deformations are allowed, since contact flexibility is simulated by means of the MATRIX27 or the COMBIN39. The rigid areas of each ring are linked to a reference node, which location coincides with the reference points in Fig. 3b. The link between these nodes and the surfaces is done by means of rigid beams. Then, the reference nodes of the rigid shell structures of both rings are linked through the corresponding MATRIX27 or COMBIN39 element. Finally, a no-separation frictionless contact is defined between the shells and rings to join them without adding circumferential stiffness to the ring. Regarding the dimensions of the contact surface, Lns is the arc length that corresponds to the angle formed by two rolling elements (2\u22c5360\u00ba / number of rollers)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003215_j.automatica.2021.109870-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003215_j.automatica.2021.109870-Figure1-1.png",
        "caption": "Fig. 1. The nonholonomic system.",
        "texts": [
            " Denote \u03be (t) = [pT (t), \u03b8 (t)]T as the state of the nonholonomic system with p(t) = [x(t), y(t)]T and \u03b8 (t) are the position and the orientation in the Cartesian coordinate frame, respectively. The nonholonomic system is described by the following kinematics \u03be\u0307 (t) = f (\u03be (t), u(t)) = [cos \u03b8 (t) 0 sin \u03b8 (t) 0 0 1 ] u(t), (1) where f (\u00b7) : Rn \u00d7 Rm is a twice continuously differentiable function with f (0, 0) = 0, u(t) is the control input with u(t) = [\u03bd(t), \u03c9(t)]T , where \u03bd(t) and \u03c9(t) are the linear velocity and the angular velocity of the vehicle, respectively. The nonholonomic system is shown in Fig. 1. In Fig. 1, \u03c5L(t) and \u03c5R(t) are the speeds of the left and right driving wheels, respectively. The upper bounds of speeds on the two driving wheels are given as |\u03c5L(t)| \u2264 a and |\u03c5R(t)| \u2264 a, where a is a known positive constant. The linear velocity and the angular velocity of the vehicle are given as \u03c5(t) = (\u03c5L(t) + \u03c5R(t))/2 and \u03c9(t) = (\u03c5R(t) \u2212 \u03c5L(t))/2\u03c1, where \u03c1 > 0 is half of the wheelbase. Therefore, u(t) belongs to the following set U = {[\u03c5(t), \u03c9(t)]T : |\u03c5(t)|/a + |\u03c9(t)|/b \u2264 1} with b = a/\u03c1. The head position \u03befh(t) is expressed as a point that lies the distance \u03c1 along the vertical bisection of the wheel axis in front of the vehicle, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure10-1.png",
        "caption": "Figure 10. Moment 2: (a) proposed model and (b) FE model.",
        "texts": [],
        "surrounding_texts": [
            "Figure 8. Finite element model in ABAQUS. will lead to a decrease in computational speed, so this article selects the number of pieces as 20. Based on the above results, the mean error of contact force along axis is 7.1% solved by the sum of the error on each slice divided by the number of slices, and the mean error of contact force of one tooth is 4.2% solved by the sum of the error at each moment divided by the number of moment sampling points, as shown in Figures 13 and 14. The comparison results verified the proposed model and the result of the proposed model well reveals the pattern of contact force distribution of beveloid gear pair. Table 3. Status of each moment. Moment Rotation angle Mesh status Moment 1 28 Two teeth in contact Moment 2 68 One tooth in contact Moment 3 108 Two teeth in contact Figure 12. Effect of the number of pieces."
        ]
    },
    {
        "image_filename": "designv11_63_0001298_j.matpr.2020.06.038-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001298_j.matpr.2020.06.038-Figure1-1.png",
        "caption": "Fig. 1. (a\u2013c). Experimental set-up of",
        "texts": [
            " and vibrational analysis of composite plates based on curve fitting method, E\u2013Glass contains one of the most significant class of fortification particularly utilized in polymer composites. It has low warm coefficient, low dielectric coefficient and high electrical opposition. Its properties rely upon added substances and restoring specialists [39]. Epoxy has great added substance properties alongside high mechanical quality, low shrinkage, artificially safe, high dissemination thickness, low gooey and better electric protection limit. What\u2019s more, it is effectively fortified with common hemp, Kenaf and E glass filaments. Fig. 1 shows the eexperimental set-up of GFRP composite for fatigue test Every composite cover comprises of and a mixture of epoxy, natural strands and e glass filaments. There are some natural fibers are utilized right now unprocessed and free from chemicals. In this study, made up of steel plate of measurement 300X300 mm is prepared. Utilizing rule of blends the different fiber weight extents are determined to accomplish covers with natural fiber and E Glass fiber. The weight extent figuring suitable measure of natural strands and gum are said something the electronic equalization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000236_s11182-019-01864-z-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000236_s11182-019-01864-z-Figure1-1.png",
        "caption": "Fig. 1. Schematic of EBAM process (\u0430) and build direction (b): 1 \u2013 build platform, 2 \u2013 AISI 321 steel substrate, 3 \u2013 grade 2 titanium prior deposit (substrate), 4 \u2013 electron beam gun, 5 \u2013 electron beam, 6 \u2013 wire feeder, 7 \u2013 wire, 8 \u2013 molten alloy puddle, 9 \u2013 re-solidified alloy.",
        "texts": [
            " The wire feedstock was conveyed from the wire-feed system, which was fixed relatively the electron beam gun as shown in Fig. 1\u0430. The electron beam parameters, the wire feeding speed and the travel speed (the ratio between the build platform motion and the beam direction) were combined so that to ensure the stability of the wire melting process and a continuous flow of the molten alloy into the molten puddle. The EBAM parameters are gathered in Table 1. The two EBAM modes differ from each other by a twofold increase in the beam current. According to the scanning procedure illustrated in Fig. 1b, three substrates are deposited onto the build platform. The specimens were prepared for metallographic examination using the automated grinding and polishing techniques. At first, the specimens were ground with abrasive papers and then polished with diamond paste. The final chemical polishing was carried out using a 2 wt.% HF\u20132 wt.% HNO3\u201396 wt.% H2O etchant in order to obtain a desired surface. The microstructure analysis was undertaken on the polished samples with an optical microscope and a scanning electron microscope equipped with an electron probe microanalyzer (EPMA)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001894_icem49940.2020.9270713-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001894_icem49940.2020.9270713-Figure3-1.png",
        "caption": "Fig. 3. Finite element models",
        "texts": [
            " The efficiency in the simulation is obtained from the torque, the iron loss, the copper loss, and the mechanical loss. The efficiency in the measurement is obtained from the input power that is obtained from the current and the voltage, the output power that is obtained from the measured torque, the mechanical loss. The efficiency is measured in a speed range going from 2,000 to 12,000 rpm and the load range going from 2 to 30 Nm. The efficiency is obtained using Maximum Torque per Amp (MTPA) as a control strategy. In this section, the finite element modeling for the efficiency map simulation is described. Fig.3 shows the 2D and 3D finite element models used for this study. In the finite element model, the circuit and the control models are included to reproduce a PWM driven current waveforms. The high harmonics components such as carrier frequency components increase the iron loss and cause the eddy current in the magnet and the winding. The iron loss is calculated as the post-processing of the finite element analysis. The hysteresis loss is calculated using the Play-Hysteron model that is a mathematical model to represent arbitrary minor loops on B-H curves of the steel sheet in the lamination core based on the actual operating condition [6]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001076_acsanm.0c00572-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001076_acsanm.0c00572-Figure1-1.png",
        "caption": "Figure 1. Schematic overview of the steps of the synthesis of EDOT-decorated NCs. (1) In the presence of acrylate monomers, anionic and cationic surfactants self-assemble into vesicles. The acrylate monomer with a photoinitiator resides in the hydrophobic region of the vesicles. The hydrophobic region may also be loaded with a pore-forming template. (2) The vesicles are extruded through a track-etch membrane and (3) the monomer in the hydrophobic region is polymerized by UV illumination. (4) The surfactant vesicle template and pore-forming template molecules are removed.",
        "texts": [
            " After cooling the mixture to room temperature, 100 mL of water is added to the reaction mixture, and the product is extracted with DCM and then washed not less than ten times or until the smell of DMSO was absent from the water wash. After removal of the solvent, the crude product was purified by column chromatography in hexane:ethyl acetate (2:1) mixture to produce a viscous yellow oil (64% yield). 1H-NMR (400 MHz, CDCl3): (ppm): 6.45 (d, 1H), 6.36 (m, 2H), 6.16 (t, 1H), 5.90 (d, 1H), 4.37 (m, 3H), 4.24 (d, 1H), 4.10 (m, 1H). Synthesis of empty and reagent-loaded EDOTDecorated Nanocapsules The steps of the synthesis of EDOT-decorated polyacrylate nanocapsules is shown in Figure 1. Page 5 of 23 ACS Paragon Plus Environment ACS Applied Nano Materials 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) Synthesis of Empty Hollow NCs: First, two separate surfactant (anionic and cationic) stock solutions were prepared by dissolving either 100 mg of SDBS (0.287 mmol) for the anionic stock solution or 100 mg of CTAT (0.219 mmol) for the cationic stock solution along with 3 mg DPA (0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure11-1.png",
        "caption": "Fig. 11. Mutual workspace area of 3-R\u0304RR manipulator.",
        "texts": [
            " For finding the mutual workspace of all the three legs, the individual workspace of each leg is translated from their base platform to a vector vi of magnitude equivalent to the radius of the mobile platform Rtop, along the direction of mobile platform orientation \u03d5\u25e6 (Fig. 10). The void circle and the reachable circle (of leg 2) are shown as red dotted circles, before the translation and black circles, after the translation by a magnitude Rtop. The reachable circles C1,C2 and C3 corresponding to each leg are translated, and the obtained mutual workspace is denoted as M1,M2 and M3 in Fig. 11. The mutual workspace area is approximated to be a circle of radius Rw, passing through the intersection points P1, P2 and P3 (Fig. 11). The intersection points are indeed the points of intersection of the medians (of base triangle) with the individual reachable circles. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S026357472000020X Downloaded from https://www.cambridge.org/core. San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at Solving by geometrical method, the radius Rw of the workspace circle is obtained. Rw = R2 d + ( lp + ld )2 + 2 \u00d7 Rd ( lp + ld ) cos ( 120 + sin\u22121 ( Rd sin ( 120 + tan\u22121 X ) cos \u03b4 )) (23) where Rd = R2 base + R2 top + 2 \u00d7 Rbase Rtop cos \u03c6 X = Rtop sin \u03c6 Rbase \u2212 Rtop cos \u03c6 lp = lp1 = lp2 = lp3; ld = ld1 = ld2 = ld3 The stiffness matrix given by (16) is used for solving the stiffness of the manipulator and is expressed as (24)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000115_j.mechmachtheory.2019.103647-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000115_j.mechmachtheory.2019.103647-Figure5-1.png",
        "caption": "Fig. 5. Graphical demonstration of the movement induced in the wheel when the swingarm moves.",
        "texts": [
            " (30) defines length L , while Eq. (31) determines the orthogonality condition of vectors r 6 \u2192 11 and r 11 \u2192 10 . x 2 + R ES cos \u03b1 \u2212 x 8 = 0 (22) y 2 + R ES sin \u03b1 \u2212 y 8 = 0 (23) x 2 + R ES cos \u03b3 \u2212 x 10 = 0 (24) y 2 + R ES sin \u03b3 \u2212 y 10 = 0 (25) x 6 + R W S cos \u03b2 \u2212 x 9 = 0 (26) y 6 + R W S sin \u03b2 \u2212 y 9 = 0 (27) x 6 + R W S cos \u03b3 \u2212 x 11 = 0 (28) y 6 + R W S sin \u03b3 \u2212 y 11 = 0 (29) (\u03b3 \u2212 \u03b1) R ES + \u221a (x 10 \u2212 x 11 ) 2 + (y 10 \u2212 y 11 ) 2 + (\u03b2 \u2212 \u03b3 ) R W S \u2212 L = 0 (30) (x 6 \u2212 x 11 )(x 11 \u2212 x 10 ) + (y 6 \u2212 y 11 )(y 11 \u2212 y 10 ) = 0 (31) Fig. 5 is used to illustrate the kinematic relationship between the two sprockets. In this Figure, it can be seen how an angular movement appears in the wheel when the swingarm moves keeping the engine sprocket locked. This movement is common in a motorcycle riding on uneven roads. This way, the relationship between the angular speeds of the engine sprocket and the wheel are not determined only by the ratio between the number of teeth of the sprockets. Therefore, the previously described angular movement has to be taken into account to properly model the whole system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure5-1.png",
        "caption": "Figure 5. Rivet bolt design model",
        "texts": [
            " To determine the parameters of the stress-strain the condition required to find solutions to ensure fatigue performance of connections, as well as for the theoretical determination of the rod retraction force, in ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 this paper. Modeled the process of rod retraction into the hole using the system of finite - of Simufact's elemental analysis. The geometry of the model is shown in Fig. 5. Geometric parameters of all elements in the model are executed in accordance with the parameters of real elements, determined by the normative and technical documentation at the Irkutsk Aviation Plant in the production of aircraft \u041c\u0421-21. These parameters can be summarized as follows. The package is represented by two fragments of 1 mm thick sheet parts made of aluminum alloy D16AT. The rivet bolt rod has a nominal diameter of 4 mm and is made of titanium alloy BT16. The model of cams is executed with idealization of the cam mechanism of the tool applied at installation of bolt rivets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000220_ecce.2019.8913300-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000220_ecce.2019.8913300-Figure1-1.png",
        "caption": "Fig. 1. Configuration of one phase of proposed Modular TC-LPMVM.",
        "texts": [
            " In order to simplify the machine configuration, a modular transverse-flux consequent-pole linear permanent magnet vernier machine (TC-LPMVM) is proposed and analyzed in this paper. The motor can be installed as any number of phases just by simply increase the number of mover modules because of its modular structure and decoupled magnetic circuit. For precisely this reason, only one mover module of the motor is analyzed. Furthermore, a Halbach PM array is applied to further increase the power factor and thrust force density. II. CONFIGURATIONS AND OPERATION PRINCIPLES The proposed modular transverse flux linear permanent magnet vernier machine is shown in Fig.1, whose primary is a split-tooth mover with a concentrated winding and a Halbach PM array secondary is applied for the purpose of increasing its thrust force density and power factor. 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 3033 Since the magnet fields of the modules of proposed TCLPMVM are separated with each other, only one mover module is adopted in this paper to explain the operation principle of the proposed machine as shown in Fig. 2. Given the symmetry of the proposed structure, only half of the model is considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000842_j.asr.2020.03.021-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000842_j.asr.2020.03.021-Figure1-1.png",
        "caption": "Fig. 1. The Earth-Centered Inertial (",
        "texts": [
            " The simulations for both the dynamic-programming-based and Lyapunov-based control in Section 4 are conducted using the coupled translational and attitude dynamics described in this section. The satellite attitude is described by a quaternion vector q defined as (Wie, 2008) q \u00bc q13 q4\u00bd T \u00bc e\u0302 sin h 2 cos h 2 T \u00f01\u00de where e\u0302 represents the Euler axis unit vector and h is the rotation angle about the Euler axis. This quaternion vector represents the rotation from the Earth-Centered Inertial coordinate system (ECI) to the body frame. In the ECI coordinate frame shown in Fig. 1, the origin is on Earth\u2019s center, the X-axis points toward the vernal equinox while the Z-axis is in the direction of Earth\u2019s rotation axis, and Y-axis completes the right-handed coordinate system. Subsequently, the kinematic equations for the satellite are written as Eq. (2). \u00f02\u00de where x is the angular velocity vector of the body frame with respect to inertial frame. Moreover, the rotational ECI) and RSW reference frames. dynamics of the satellite as a rigid body are described by the general form of Euler\u2019s rotation equation (Wie, 2008) _x \u00bc J 1 x Jx\u00feMB \u00f03\u00de where J3 3 denotes the inertia matrix and MB is the total external torque vector generated by the thrusters. It should be noted that the other perturbation effects are negligible in comparison to the thrusters\u2019 forces and moments. The relative position between the chaser and its target q \u00bc dx dy dz\u00bd T is expressed in the RSW frame shown in Fig. 1. In this co-moving frame, the x-axis is aligned with the position vector R from the Earth center to the target satellite, z-axis is normal to the orbital plane in the direction of the orbital angular momentum, and y-axis completes the right-handed coordinate system. The translational dynamics are given by the linearized equations of relative motion in orbit (Curtis, 2013) \u20acq\u00bc f q; _q; t\u00f0 \u00de\u00bc 2l R3\u00fe h R4 dx 2 V :R\u00f0 \u00deh R4 dy\u00fe 2h R2d _y h R4 l R3 dy 2 V :R\u00f0 \u00deh R4 dx\u00fe 2h R2d _x l R3 dz 2 6664 3 7775\u00fe 1 m F R x F R y F R z 2 64 3 75 \u00f04\u00de where R and V represent the position and velocity vectors of the target orbit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001656_etfa46521.2020.9211891-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001656_etfa46521.2020.9211891-Figure2-1.png",
        "caption": "Fig. 2: Active contact flange of FerRobotics",
        "texts": [
            " 1. Research on this specific problem seems not to have been addressed in the literature. The simpler problem of automatically connecting to a charging station has been addressed in context of electric vehicles [5], [6]. This paper presents a method for the automated plugging of an electric connector for the purpose of functional quality testing. The mathematical formulation, the algorithmic steps as well its implementation in a product testing facility in a production environment are presented. Fig. 2 shows the product to be tested. This is the Active Contact Flange (ACF), manufactured by the company FerRobotics [7]. The ACF is a mechatronic actuator as well as sensor element and a robot equipped with this product is able 978-1-7281-8956-7/20/$31.00 \u00a92020 IEEE 1465 Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 02,2020 at 11:44:48 UTC from IEEE Xplore. Restrictions apply. to apply a consistent contact force on a surface. Due to the ACF\u2018s one degree of freedom, it is not required to adapt the robot path because it can compensate for surface tolerances up to 100 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001528_0954407020951318-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001528_0954407020951318-Figure1-1.png",
        "caption": "Figure 1. DMF structure.",
        "texts": [
            " Firstly, the DMF model is established and verified by the bench test. Secondly, the smooth function is used to fit the segmented points in the DMF model, and the influence of fitting factor b on the smooth fitting effect is studied. Finally, a 3DOF driveline system model including DMF is established, and the effectiveness of the smoothing function is analyzed by simulation, and the performance differences between different smoothing functions are studied The basic structure and working principle of DMF The structure of the DMF is shown in Figure 1. DMF is composed of first mass, secondmass, startingring, force transmission plate, and long arc-spring. The starting ring is connected to the first mass by an interference fit. The first mass is bolted to the end flange of the engine crankshaft. The flange is connected to the second mass by a rivet, and the second mass is fixedly connected with the clutch assembly by bolts. The first mass assembly and the second mass assembly are connected by a low-rigidity arc-spring and are rotatable relative to each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002511_012004-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002511_012004-Figure10-1.png",
        "caption": "Figure 10. Maximum stress results for different shapes.",
        "texts": [
            " Besides, in addition to the common round whole profile, other profiles are also cut, such as alternating grooves, large and small round holes (Figure 9). However, the number of holes must be an even number so as to avoid the eccentricity when the gears are operating. RCMEManuE 2020 IOP Conf. Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10.1088/1757-899X/1109/1/012004 Based on the element removal results in Section 2.1 and practical experience, the researcher cuts the 3D gear model at the disk part with different shapes and rechecks the maximum stress (Figure 10). RCMEManuE 2020 IOP Conf. Series: Materials Science and Engineering 1109 (2021) 012004 IOP Publishing doi:10.1088/1757-899X/1109/1/012004 To test the teeth strength, we use ISO 6336 [2]. Allowable contact stress is determined by formula (2) to obtain a value of 608.8 N/mm2 (MPa). The volume percent reduction, compared to the original model according to the different shapes, is shown in Table 3. As mentioned above, during the operation of the gear, there is only one bearing tooth at a time when the gear is engaged"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002170_s00170-020-06399-z-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002170_s00170-020-06399-z-Figure6-1.png",
        "caption": "Fig. 6 a Predicted topography of top surface (relative to highest point) at conclusion of SLM process to build 4 mm \u00d7 4 mm 316L stainless steel block having 200-\u03bcm part height obtained using four 50-\u03bcm layers. b Color-contoured region on top surface that fails to meet the assigned 2-\u03bcm conformance limit (non color-contoured region meets 2-\u03bcm conformance criterion)",
        "texts": [
            " Although 316L is considered in this work (not Inconel 625), the volumetric heat flux due to the laser incidence assumes the same double ellipsoid parameters; this is justified in that the intent of this work is to assess the viability of correcting typical SLM distortion patterns using in situ LSP, rather than a rigorous predictive modeling of the actual distortion field itself. Following simulation with the described SLM model using the geometry and scan parameters in Fig. 4, as well as Tables 1 and 2, the predicted topography of the top surface is seen in Fig. 6a (based on 6561 upper surface nodes). It is noted that the topography exhibits a convex deflection similar to that described by Li et al. [15]. From surface roughness metric Rt (range), it is evident that a \u223c 9-\u03bcm vertical distortion exists between the highest and lowest points. Figure 6b shows the region of the top surface that falls outside an assigned 2-\u03bcm conformance limit (used in this work for demonstration purposes). The colorcontoured region, which shows \u223c 7 \u03bcm of variation, is to be treated with laser shock peen forming, as described next in Sections 3 and 4, to preferentially form the surface so as to strive for maximum conformance. Surface roughness metrics, such as Ra (arithmetic mean) and Rq (root mean square), are also quantified and discussed to further assess the merits and effectiveness of LSP forming treatments",
            " 10a shows the plasma pressure temporal profile for LSP shots having 754-MPa peak pressure (corresponding to 21 MW cm\u22122 laser peak power density obtained with an averaged incident laser pulse energy of \u223c 29 mJ), whereas Fig. 10b shows the plasma pressure temporal profile for LSP shots with 904.8-MPa peak pressure based on 50% increase in laser peak power density to 31.5 MW cm\u22122 obtained with an averaged incident laser pulse energy of \u223c 42 mJ. For convenient reference in analyzing results of LSP forming, Fig. 6a and b are repeated as Fig. 11a, b, to show the vertical displacement (distortion) contour maps, relative to the highest point, on the top surface of the 316L part following SLM of the four 50-\u03bcm layers. Again, the noncontoured (white) region in Fig. 11b represents the top surface area conforming to the assigned 2-\u03bcm tolerance criterion following SLM. The relative area of conformance is calculated to be 44.13%. The non-conforming region has a vertical distortion range of \u223c 7 \u03bcm and will thus be subject to LSP forming treatment (in practice LSP forming would occur prior to any further SLM layer deposition)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002374_j.matpr.2020.11.841-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002374_j.matpr.2020.11.841-Figure4-1.png",
        "caption": "Fig. 4. Model of car seat and the deformation car seat.",
        "texts": [
            " Vibration test fixtures are required to allow mounting of the test specimen to the fix as well as to allow for testing in all the three orthogonal directions. The design of vibration test fixtures is critical to avoid errors in equipment test response due to any resonances of the impact hammer, table and the fixture itself [12]. Ideally the laboratory mounting should replicate the physical conditions observed in service such as the stiffness, mass and the consequent resonant responses of the actual service installation. The model of car seat and the deformation car seat is shown in Fig. 4. Fig. 5 shows the Von moises stress and displacement in the car sheet. The deformation present in car sheet surface with frequency analysis is shown in Fig. 6. The details of stress versus displacement of ABS, polyurethane and Nylon materials observed by Modal (without bush) analysis are given in Tables 1\u20133. Cost reduction. Less waste. Reduce production time. An enhanced competitive advantage. Reduce errors. Confidentiality. Production on demand. Surface texture is generally too rough. Materials have low heat deflection temperatures Materials generally have low strengths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001123_5.0009667-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001123_5.0009667-Figure11-1.png",
        "caption": "FIG. 11. Two spherical caps at r = R of angle \u03b1, whose centers are separated by the polar angle \u03b2 and azimuthal angle \u03c9.",
        "texts": [
            " Finally, we shall employ mean field arguments to compute the total interaction energy U inter for N domains, which implicitly neglects structural details beyond their average density. It is expected that these simplifications will give qualitative, though not quantitative, insight into the effect of the surface curvature on domain\u2013domain interactions. We now delve into the mathematical details of our model for dipolar interactions on a sphere. The natural starting point is to consider a pair of equal-sized, spherical-cap domains of cap angle \u03b1 whose centers are separated by a polar angle \u03b2 and azimuthal angle\u03c9 (Fig. 11). Using the addition theorem for Legendre functions (Hobson,109 p. 143), Pl(cos \u03b8\u2032) = Pl[cos\u03b2 cos \u03b8 + sin\u03b2 sin \u03b8 cos (\u03d5 \u2212 \u03c9)] = l \u2211 m=\u2212l eim(\u03d5\u2212\u03c9) (l \u2212m)! (l + m)! Pm l (cos \u03b8)Pm l (cos\u03b2), (49) we obtain the pairwise interaction energy U(\u03bc)12 = \u2212 \u03c0\u03bc2R sin2 \u03b1 \u03b5 \u221e \u2211 l=1 [P1 l (cos\u03b1)]2Pl(cos\u03b2) l(l + 1) , (50) where Pl \u2261 P0 l is the Legendre polynomial of degree l. The sum on the right-hand side of (50) is convergent for \u0394/R = 0 and may be expanded for sufficiently large separations as U(\u03bc)12 = \u03c0\u03bc2R sin4 \u03b1 8\u03b5 ( 3 \u2212 cos\u03b2 (2 \u2212 2 cos\u03b2)3/2 +\u22ef), (51) where the terms inside the brackets are associated with higher-order multipole moments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001706_s12555-020-0031-7-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001706_s12555-020-0031-7-Figure3-1.png",
        "caption": "Fig. 3. 2R planar manipulator, a two-link robot with revolute joints.",
        "texts": [
            " Note that (2) provides the dynamics of the ith link. It should also be observed that the dynamics of the ith link has parameters (such as mass, and link length), states (\u03b8 and \u03b8\u0307 ), and the derivative of states, corresponding to other links. This captures the coupling of dynamics in a subtle manner. This is precisely the \u2018distributed\u2019 nature of manipulator dynamics as discussed in Section 2. Consider the dynamics of a two-link planar manipulator with revolute joints having mass concentrated at the end of the links as shown in Fig. 3.( \u03c41 \u03c42 ) =  m2l2 2 +2m2l1l2c2 +(m1 +m2)l2 1 m2l2 2 +m2l1l2c2 m2l2 2 +m2l1l2c2 m2l2 2 (\u03b8\u03081 \u03b8\u03082 ) + ( \u2212m2l1l2s2\u03b8\u0307 2 2 \u22122m2l1l2s2\u03b8\u03071\u03b8\u03072 m2l1l2s2\u03b8\u0307 2 1 ) + ( m2l2gc12 +(m1 +m2)l1gc1 m2l2gc12 ) . (3) Equation (3) can be written as \u03c41 =\u03c4\u03032 + \u03c4\u03031; \u03c42 = \u03c4\u03032, (4) where \u03c4\u03031 =(m2l1l2c2 +(m1 +m2)l2 1)\u03b8\u03081 +m2l1l2c2\u03b8\u03082 \u2212m2l1l2s2\u03b8\u0307 2 1 \u2212m2l1l2s2\u03b8\u0307 2 2 \u22122m2l1l2s2\u03b8\u03071\u03b8\u03072 +(m1 +m2)l1gc1, (5) \u03c4\u03032 =(m2l2 2 +m2l1l2c2)\u03b8\u03081 +m2l2 2 \u03b8\u03082 +m2l1l2s2\u03b8\u0307 2 1 +m2l2gc12. (6) We may write this as \u03c4\u0303 = M\u0303(\u03b8))\u03b8\u0308 +V\u0303 (\u03b8 , \u03b8\u0307)+ G\u0303(\u03b8)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003238_s12206-021-0829-0-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003238_s12206-021-0829-0-Figure4-1.png",
        "caption": "Fig. 4. Gear and bearing temperature contours.",
        "texts": [
            " 3) Energy-conservation equation ( ) ( ) ( ) ( ) T p p p S c T uT vT wT t x y z k T k T k T x c x y c y z z \u03c1 \u03c1 \u03c1 \u03c1\u2202 \u2202 \u2202 \u2202 + + + = \u2202 \u2202 \u2202 \u2202 \u239b \u239e \u239b \u239e \u239b \u239e\u2202 \u2202 \u2202 \u2202 \u2202 \u2202 + + +\u239c \u239f \u239c \u239f \u239c \u239f\u239c \u239f \u239c \u239f \u239c \u239f\u2202 \u2202 \u2202 \u2202 \u2202 \u2202\u239d \u23a0 \u239d \u23a0 \u239d \u23a0 (13) where T is the fluid element temperature, K; k is the heat conductivity of the fluid, W/(m\u00b7K); cp is the specific heat capacity of the constant pressure fluid, J/(kg\u00b7K); and ST is the viscous dissipation term, W. If the residual curve of the parameters in the massconservation, momentum-conservation, and energy-conservation equations tends to be stable, then the simulation has converged. Fig. 4 shows the temperature field contours of the sun gear, planetary gear, ring gear, and planetary gear bearing when the diameter of the planetary gear opening is \u042420 and the number of openings is six (denoted as \u042420\u00d76). The temperature of the inner raceway of the planetary gear bearing is lower than that of the outer raceway due to the influences of the under-ring lubrication and heat generation of the planetary gear. Fig. 5 shows the contours of the convective heat transfer coefficients of key moving components, such as the sun and planetary gears, when the opening parameter of the planetary gear is \u042420\u00d76, and the maximum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001249_s00170-020-05744-6-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001249_s00170-020-05744-6-Figure5-1.png",
        "caption": "Fig. 5 Exploded view of actuator assembly",
        "texts": [
            " The actuator design consists of four main types of parts as follows: a top layer of chitosan films, rigid members that compose the body of the actuator, 3D printed nodal joints, and mechanical components for enabling rotation. The movement mechanism for the actuator is based on chitosan\u2019s ability to change its mechanical properties with water content, as was demonstrated by the mechanical testing results in Section 4.1.4, as well as a degree of freedom that is stored in the nodal joints. 3D-printing enables the customization of the nodal joints for accommodating metal parts that induce rotational freedom as shown in Fig. 5. The joint design consists of two parts, a male and female connection, that when fitted together act as points of convergence for the various members of the actuator body. The male half features an extrusion with a cavity for a ball bearing, while the female half envelops the extrusion on either side with a tight fit. Both halves feature a hole through their central axis for a dowel pin to be inserted which acts as the rotational member and connects the two halves. Frictionless rotation is achieved by ensuring there is no friction in the joint and all surfaces have a smooth finish, which is provided by giving proper sliding tolerance to the 3D printed parts following design for additive manufacturing (DFAM) standards"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure6.7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure6.7-1.png",
        "caption": "Fig. 6.7 Schematic representation of a\u00a0gear produced by a DMLS system. (a\u2013c) The component is formed layer by layer",
        "texts": [
            " Unlike a powder feed system, where the particles have the purpose of melting by using a higher-energy concentration in their union, sintering provides only the energy necessary to cause atomic diffusion among the particles that are in contact with each other, forming metallurgical bonds and providing mechanical stability to the whole. 6 Additive Manufacturing There is a wide variety of commercial elements and alloys available for the manufacture of different architectures from a design generated on a computer. During the DMLS process, a laser is used to sinter the first layer of metallic powder placed on the base plate, in order to generate the support structure (Fig.\u00a06.6). Then the base plate moves down allowing a new layer of metallic powder to be deposited (Fig.\u00a06.7). The sintering process of the subsequent layer follows the pattern generated by the computer design, successively until the manufacturing process is completed. Once the process is finished, the loose powder is removed. DMLS is a technology that relies on the use of metal powders for their later projection at a given pressure and speed. Most of the powders used in additive manufacturing are produced by atomization [16] (Fig.\u00a06.8). The water atomization method is the most widely used for the production of metallic powders due to its low cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001450_1464420720948541-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001450_1464420720948541-Figure7-1.png",
        "caption": "Figure 7. Deviation field between the scanned heat-treated samples and the theoretical 3D cylindrical shapes of the single-filed (a) and multi-filed (b) specimens.",
        "texts": [
            " Samples of virgin powder from the same (Figure 6 (e)) and different (Figure 6(f)) supplier of the used powder are shown. No significant changes are identified between re-used and virgin powders from the same supplier. Distortion analysis was conducted by digitally overlapping the scanned geometries with the theoretical cylindrical shapes. Scanning was performed before Theoretical density, q (g/cm3) 7.78 0.01 7.83 0.0125 Apparent density, q0(g/cm3) 4.762 4.45;26 4.8127 Internal friction coefficient, k (Pa) 93.33 196 325,28 specimens removal for as-built and post-heat treatment stages. Figure 7 shows the deviation to the 3D theoretical geometries of the heat-treated specimens (before removal from the build plate). It is visible that for both single and multi-field specimens, maximum deviations to the theoretical geometry range in a 1 mm tolerance [\u20130.5; \u00fe0.5 mm]. In order to have a more accurate measurement of each specimen and to identify eventual differences among them, the contour of their longitudinal section (refer to the signalled Sa section in Figure 7(a)) was subject of analysis. In that section, Sa, for each specimen, the maximum geometric deviation, dmax, was taken into consideration for calculating the geometric deviation ratio, dr, as displayed in equation (1), where L corresponds to the specimens\u2019 length. An example of the considered contour for D1 (multi-field) specimen is shown in Figure 8. In addition, the deviation of each cylinder can be compared in Figure 9 for asbuilt and heat-treated conditions dr \u00bc dmax L (1) A clear distinction between single and multi-field specimens, regarding their deviation with respect to the theoretical shape, can be observed in Figure 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000439_s40684-019-00177-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000439_s40684-019-00177-3-Figure1-1.png",
        "caption": "Fig. 1 Structure laminated by the FDM",
        "texts": [
            " The experimental and analytical investigation showed that the perimeter has much stronger effect on the FDM structures than the core lattice except for small raster angle conditions. Because the joints of core part behave like a hinge, and the hinge effect increases as raster angle increases. To verify this joint effect, additional experiments were conducted with special specimens having no side layer. The results of this work induce that the side layer should be well designed for FDM manufactured structures to resist external forces. Figure\u00a01 shows the main components of a structure made by FDM process. The definitions of process terms are explained in Table\u00a01. As shown in Fig.\u00a01a, an FDM-manufactured structure has a core part, side layer (or perimeter), and top and bottom layers. The role of the bottom layer is to provide adhesion between the additive manufactured structure and the bed of the 3D printing machine. The role of the top layer is in the finishing process of additive manufacturing. The side part maintains the shape of the whole structure, and the core part is laminated to fill the space inside the side part after the side layer is built. The infill density parameter determines the volume fraction. In many engineering cases, to reduce the process time and energy, the infill density is set at 20\u201330%, and the core part is then built automatically as a truss lattice. From the viewpoint of structural analysis, since the side and core parts are the main components that withstand external forces, these two components should be well designed. For this reason, this work focused on these two parts, as shown in Fig.\u00a01b. Figure\u00a02 shows examples of the core lattice with respect to infill density and raster angle. The target shape of this study was the ASTM D638 specimen with a 5\u00a0mm thickness for the tensile test. Detailed specimen geometry is shown in Fig.\u00a03. For the experimental study, the tensile test specimens were made using a Davinci Pro 1.0 machine; the fixed settings are summarized in Table\u00a02. The main variables of the FDM process were side layer thickness, infill density, and raster angle. The Davinci Pro 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure15.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure15.6-1.png",
        "caption": "Fig. 15.6 Orientation of parts on 3D printer build plate to avoid more support structures and get good surface finish",
        "texts": [
            " \u2022 All 3D components were modeled in Autodesk Fusion 360, and Ultimaker Cura was used as slicing software for generating GCODE for 3D printer. 3D printing of parts was carried out in Anycubic Kossel 3D printer (Liner plus) with build volume of 230 \u00d7 230 \u00d7 300 mm. Components were designed in a way that can be easily printed in most common 3D printers with build area greater than or equal to 200\u00d7 200 mm. 3D printing was carried out using poly lactic acid (PLA) filament. The whole body of spin coater was divided into nine components, and only bio-inspired components are shown in Fig. 15.6, i.e., (a) leg, (b) lid, (c) vibration isolation shoe, (d) electronics enclosure, (e) spin drum, (f) display and control panel and vacuum less chuck, chuck holder, collar are not shown. Unlike bio-mimicry where one has to exactly copy the structural design in nature based on size and morphology. Bio-inspired design gives flexibility to change the design and further enhance it based on aesthetic or functional needs. Before designing the model of the proposed product, designer has to consider its overall volume or size, which further depends on factors like need for portability, electronics encased within it and its functional aspects. In our work, we found that size of the electronics enclosure (Fig. 15.6d) would decide the final volume of the spin coater as it accommodates important electronic components like brushless DC motor, driver and control circuitry. Due to limited build area of 230 \u00d7 230 mm, dimensions of electronics enclosure were restricted to less than 200 \u00d7 200 mm. Considering all these attributes, itwas decided to design electronics enclosurewith a toroidal space to hold circuitry,wiring and apost at the center to hold brushless direct current (BLDC) motor. Also size of the overall parts which are too big to 3D print can be designed symmetrically along with bio-inspired aspects and can be divided into small 3D printable parts based on symmetry. Here, the supporting leg of spin coater inspired by tortoise was found to be too big to be directly 3D printed; hence, it was 3D printed after dividing into four symmetric parts as shown in Fig. 15.6a. Later, instant glue is used to stick four parts together. Likewise, packaging foam was cut into cuboids shape of desired dimensions and glued to base of vibration isolation shoe (Fig. 15.6c). All the components are modeled for minimal to no need for support structure for 3D printing allowing good quality of print and minimal wastage of printing material. After 3D printing, some parts may require post-processing like sanding for smoothening surface and drilling for screw fastening. All these factors which was considered for additive manufacturing of parts for spin coater can be easily extended for any part design for products. Final assembled spin coater is as shown in Fig. 15.7. \u2022 Testing of spin coater is done by evaluating two aspects, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure4-1.png",
        "caption": "Fig. 4. Top view of the workspace of RR configuration.",
        "texts": [
            " Although the links are made inclined to the horizontal plane, the rotary (or prismatic) joints axes remain perpendicular (or parallel as the case may be) to the plane of the base platform, which retains the planar motion of the end effector. New mathematical models are required for studying the influence of the inclination angle in relation to the following parameters, such as workspace, MOI, stiffness, static forces and the moving mass of the manipulator. The workspace of the individual serial configuration (leg i) has been obtained geometrically. The non-planar links and its workspace are shown in blue color (Fig. 4). To have a meaningful comparison of the PKMs, the workspace area should be same for both the manipulators, and hence the same link lengths are used. If the link lengths are taken as l1 and l2 for the conventional planar configuration, then the projected orthographic link lengths for NPL should also be taken as l1 and l2. It is because the projected link lengths are only applicable for computing the workspace of NPL. These lengths are divided by cos\u03b4 to get the original isometric link length. https://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003143_acc50511.2021.9483293-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003143_acc50511.2021.9483293-Figure1-1.png",
        "caption": "Fig. 1. Euler angles [11] : positive values are indicated from FS (green) to FR (red)",
        "texts": [
            " The kinematics of the attitude is R\u0307 = R\u2126\u0302, where \u2126 \u2208 R3 is the angular velocity of the body resolved in FB . \u2022 Right wing: Let FR = {rx, ry, rz} be the frame fixed to the right wing at its root. And let FS = {sx, sy, sz} be the stroke frame obtained by translating the origin of FB to the center of wing roots, and rotating it about by by a fixed angle \u03b2 \u2208 [\u2212\u03c0, \u03c0). Let \u00b5R \u2208 R3 be the fixed vector from the origin of FB to that of FR. The attitude of the right wing relative to FS , namely QR \u2208 SO(3) is described by 1\u2013 3\u20132 Euler angles (\u03c6R(t), \u03c8R(t), \u03b8R(t)) (Figure 1) as QR = exp(\u03b2e\u03022) exp(\u03c6Re\u03021) exp(\u2212\u03c8Re\u03023) exp(\u03b8Re\u03022), and its time-derivative is Q\u0307R = QR\u2126\u0302R for \u2126R \u2208 R3. \u2022 Left Wing: Similarly, for the left wing, QL = exp(\u03b2e\u03022) exp(\u2212\u03c6Le\u03021) exp(\u03c8Le\u03023) exp(\u03b8Le\u03022), with the set of Euler-angles (\u03c6L(t), \u03c8L(t), \u03b8L(t)), and Q\u0307L = QL\u2126\u0302L for \u2126L \u2208 R3. \u2022 Abdomen: The abdomen is considered as a rigid body attached to the body via a spherical joint. The frame fixed to the abdomen is FA = {ax,ay,az}, and its attitude relative to the body is denoted by QA \u2208 SO(3) with Q\u0307A = QA\u2126\u0302A for \u2126A \u2208 R3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002239_s11771-021-4591-3-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002239_s11771-021-4591-3-Figure3-1.png",
        "caption": "Figure 3 Dynamic model of planetary gear train",
        "texts": [
            " (2021) 28: 126\u2212139 128 rotation speeds are np and nr, respectively. P is the pitch point, and K denotes any possible meshing point during the transmission process. For high-speed and heavy-duty gear systems, the damping ratio and inertial forces have significant effects on the transmission characteristics, and thus, dynamic analysis is necessary. The dynamic model of the planetary train was established accounting for the excitation of the time-varying meshing stiffness [23\u221225]. The dynamic model is shown in Figure 3. For the inner gear, there is r r rp rp rp rp br p( )J c k r T       (1) where Jr represents the moment of inertia of the inner gear, \u03b8r is the torsional angular displacement induced by the vibration, krp and crp are the meshing stiffness and damping between the planetary gear and the inner gear, respectively, \u03b4rp is the displacement of the planetary gear and the inner gear at point K along the line of action (LOA), and Tp is the input torque. For the planetary gear, there is p p sp sp sp sp bp rp rp rp rp bp( ) ( ) 0J c k r c k r             (2) where Jp represents the moment of inertia of the planetary gear, \u03b8p is the torsional angular displacement caused by the vibration, ksp and csp are the meshing stiffness and damping between the planetary gear and sun gear, respectively, and \u03b4sp is the displacement of the planetary gear and the sun gear at point K along the line of action"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001968_me49197.2020.9286694-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001968_me49197.2020.9286694-Figure3-1.png",
        "caption": "Fig. 3. UAV control, a thicker line means an increased velocity",
        "texts": [
            " Recent studies have shown that the first modes of vibration related to structural deformation of the octocopter appear above 40 Hz [12]. This UAV, as most of multi-rotor UAVs, is controlled acting on propellers angular velocities. Increasing or decreasing the 8 propellers angular velocities together changes the global lift force and generates the vertical motion. Increasing the angular velocity of propellers 1, 2, 3, 4 and decreasing of the same amount the angular velocities of propellers 5, 6, 7, 8 does not change the global lift force, but generates a torque about axis that controls the roll rotation (\u03d5), see Fig. 3. A similar operation carried out on two different groups of propellers (3, 4, 5, 6 and 7, 8, 1, 2) generates a torque about axis that controls the pitch rotation (\u03b8), see Fig. 3. Increasing the angular velocities of the 4 propellers rotating clockwise and decreasing of the same amount the angular velocity of the 4 propellers rotating counter-clockwise generates a torque about axis that controls the yaw rotation (\u03c8). It is worth highlighting that this UAV is under-actuated, because there is no possibility of directly generating thrusts in the and directions. Actually, the roll and pitch rotations alter the direction of the thrust with respect to the inertial reference frame and make possible motions along the and directions",
            " It is equal to zero, if the angular velocities of the propellers rotating clockwise are equal to the angular velocities of the propellers rotating counter-clockwise. The term is the global thrust force that acts along . In hovering flight it is almost aligned to and has small components along and . The terms , , are the torques about the body axes generated by the propellers. The term is the barycentric moment of inertia of propellers around their axis, and and are the lift and drag coefficients of propellers. Angles \u03b2 are the angles between and the rotor arm (see Fig. 3), and (\u03b2 ) and (\u03b2 ) are their sine and cosine. A control system is needed to set the inputs , , , in order to obtain the desired vertical position and attitude of the UAV. Many control strategies have been proposed to control the attitude of UAVs, see for example the review paper [13]. In this research, the control problem is simplified, since small oscillations about the hovering configuration are considered. In this case, the coupling terms between roll, pitch, and yaw equations become negligible and a set of independent PID controls [14] is used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002139_icarcv50220.2020.9305333-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002139_icarcv50220.2020.9305333-Figure1-1.png",
        "caption": "Fig. 1: Fleet of multi-robots in agriculture (a) Master-slaves convoy, (b) Convoy on pre-assigned trajectories",
        "texts": [
            " To fulfill some of these challenges, the potential of mobile robotics in terms of accuracy, repeatability and work capacity is increasingly put forward as a means to increase field efficiency and develop new environmentally friendly practices, as well as to relieve human operator from unhealthy operations [2], [3]. In the last few years, combined interventions of several agricultural robots in the same field have received increased interest to preserve soils from compaction and improve flexibility with the deployment of fleet of more or less numerous light robots [4], [5]. Immaterial towing of farm vehicles, as initially developed for convoy of trucks on highway [6] and usually called master/slaves convoy [7], is the main adopted strategy in the literature, see the illustration on Figure 1a: the master vehicle is manually driven by a human operator and the slave vehicles have to accurately follow its trajectory \u0393master with respect to a desired curvilinear distance si, but also with respect to a lateral deviation yi depending on the working width of the vehicles, see [5], [8]. The main advantage of this approach is to not require a preliminary step of trajectory planning as the trajectory \u0393master is created on-line, typically from the positions transmitted by radio link to the slave vehicles with a RTK GPS embedded on the master vehicle",
            " firstname.lastname@inrae.fr be feasible for the followers. In practice, such an approach involving shifted following vehicles is difficult to control for the human operator, especially during turns and maneuvers, but also in presence of obstacles and when the fleet is composed of a high number of vehicles. Moreover, farm fields can have complex irregular shapes leading to difficult on-line control of a fleet of robots in such environments [11]. The convoy on pre-assigned trajectories, illustrated on Figure 1b, is another strategy of multi-robots system: each vehicle has for mission to accurately follow its own previously planned trajectory \u0393A,B,C,...,Z and operates independently, eventually completed by a collision avoidance system or by imposing a safety inter-vehicle distance [9], [10]. The main advantage is that all the trajectories can be previously planned and adapted to both the field geometry, the known obstacles within the field, the steering and speed capacities of the vehicles as well as the agricultural tasks to be performed",
            " Downloaded on May 27,2021 at 15:28:54 UTC from IEEE Xplore. Restrictions apply. trajectories based on adaptive clothoids are generated and the trajectories of U-turn for the fleet of vehicles are built. The main driving direction in the field is determined from a bruteforce algorithm. Finally, experimental fields with different shapes are processed to highlight the capabilities of the multirobots trajectory planner proposed. To address the trajectory planning of multi-robots on preassigned parallel trajectories as depicted on Figure 1b, the method proposed in this paper is summarized in Figure 3. A virtual robot, positioned in the middle of the fleet and gathering the steering and speed constraints of the n vehicles, is first considered. A trajectory generation algorithm suited for this virtual robot is then developed through the adaptation of the shape of clothoids with respect to the motion constraints of the robot. The U-turn maneuvers for the fleet in headland are next considered. A first strategy consists to maintain the geometrical layout of the fleet during the maneuver"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure59.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure59.3-1.png",
        "caption": "Fig. 59.3 Computation of major parameters for consolidation of Concept 3",
        "texts": [
            " Turning to the problems of our focus, the robot can reach narrow spaces and under obstacles but must be placed manually over obstacles, if needed. In its manually controlled version, the user detects the area to be cleaned and manoeuvres the robot. The extendable tube allows cleaning of narrow areas. Detachability ensures ease of use whenever needed. The use of suction as well as a rotary brush ensures cleaning of the surfaces. Each concept is consolidated further by identifying themain components required for its embodiment. For example, the calculations and estimations of motor power, etc., for Concept 3 are given below, see Fig. 59.3. The following are taken as the governing equations for Concept 3: F = Fs + Ft + Fr (59.1) Ft = (Mt )(g)(\u03bc) (59.2) Fs = (p)(At ) (59.3) Fr = (Mr )(\u03bc) (59.4) T = (F)(r) (59.5) P = (T )(\u03c9) = (Vr )(F) (59.6) Typical suction power for robotic vacuum cleaners starts at 8 AW. With some factor of safety for losses due to tube-length, leakage, etc., using an indicative vacuum pump (suction power = 14 AW; flow rate = 1.5 LPM; suction area = 3 cm2; suction pressure = 5.5 bar) for calculation gives a maximum permissible robot velocity of 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure5-1.png",
        "caption": "Fig. 5 Static stiffness analysis setup illustration",
        "texts": [
            " (6)D = pE 1 \u2212 2 \u23a1 \u23a2\u23a2\u23a3 1 0 1 0 0 0 1\u2212 2 \u23a4 \u23a5\u23a5\u23a6 (7)K(x) = n\u2211 e=1 xp e K0 e (8)Ei = Ei ( xi ) = x p i E0, xi \u2208 [0, 1] (9)Ei = Ei ( xi ) = Emin + x p i ( E0 \u2212 Emin ) , xi \u2208 [0, 1] (10)E(x) = f ( x)E0 (11)Vi = \u222bV0 i (x)dV 1 3 Static stiffness is calculated by applying unit loads to the engine mounting positions in all degrees of freedom. The displacement is then extracted from the loaded point in all degrees of freedom, and the unit load is divided by that displacement to obtain the stiffness [18]. The boundary conditions for the cross member are set up to replicate the attachment point stiffness of a body in white and thus provide as realistic of a measure of stiffness as possible. A static stiffness analysis setup illustration is shown in Fig.\u00a05. The relationship between stiffness and displacement can be written as follows [17, 19, 20]: where u is the nodal displacement vector and F is the nodal forces vector. The nodal displacement is solved as follows [17, 19, 20]: The dynamic stiffness calculation is basically similar to the static stiffness calculation, with the only difference being the specific implementation form of the load. For dynamic stiffness, the unit load is applied as a dynamic excitation over a defined frequency range, which can be shown in a displacement or stiffness versus frequency plot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure12-1.png",
        "caption": "Fig. 12 Push-pulling the hook model and the modeling results",
        "texts": [
            ") There are in total eight case studies, which are based on six real-world mechanical parts obtained from the GrabCAD part library (https:// grabc ad. com/ libra ry). Case study 1 considered push-pulling an axis support part model where no critical points could be detected during the push\u2013pull edit (Fig.\u00a010). Case study 2 involved a connecting rod part model with one critical point in the push\u2013pull edit (Fig.\u00a011). Case study 3 analyzed push-pulling a hook part model with one critical point in the push\u2013pull edit (Fig.\u00a012). The varied modeling results in these three case studies will be used to show the essential role GTI detection plays in attaining robust push\u2013pull with G1 connections, as well as the effectiveness of the proposed GTI detection method. Case studies 4\u20137 considered three comprehensive modeling examples where various push\u2013pull edits were applied (Figs.\u00a013, 14, 15, 16); they will be used to show the effectiveness of the proposed method as a whole. Case study 8 (Fig.\u00a017) is intended to show some important limitations of the proposed method",
            " In case study 2, similar faces to those in case study 1 were push-pulled, but an invalid modeling result was generated in Siemens NX, and a valid yet unpredictable result (as indicated by the red circle) was given by SpaceClaim (Fig.\u00a011). (Siemens NX colors boundary faces in red whenever there is a model update failure, as shown by the circled face in Fig.\u00a011.) The major difference between case studies 1 and 2 is that the latter involves a critical point of GTI. It can thus be concluded that crossing critical points could cause model update failures. In case study 3 (Fig.\u00a012), there was also one critical point of GTI, and the GTI configurations were almost the same as those in case study 2. Despite the similarity, Siemens NX failed in case study 2 and succeeded in case study 3, thereby leading to the conclusion that the failure in case 1 3 study 2 was likely not due to GTI resolution but GTI detection. The failure was not likely due to numerical instability [23] either, because the geometric configurations in the two cases are also very similar. As such, the significance of GTI detection in attaining robust push\u2013pull with G1 connections can be partly confirmed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure12-1.png",
        "caption": "Fig. 12 The schematic diagram",
        "texts": [
            "\u00a011) must complete three working requirements: horizontal folding, vertical folding and reset. (1) Horizontal folding, as shown in Fig.\u00a011a, horizontally pushing the left side of the single-layer cardboard and rotating the second side of the left side along the second crease to vertical position. (2) Vertical folding, as shown in Fig.\u00a011b, completing the rotation of the first surface around the first crease and coinciding with the second surface. (3) Reset, as shown in Fig.\u00a011c, the mechanism returns to its initial position. Figure\u00a012 shows the schematic diagram of the planar double-folded metamorphic mechanism and its composition principle. According to the above requirements, force metamorphism is adopted at kinematic joint D, that is, when the mechanism is in configuration 1, the spring force is set at kinematic joint D, so that the relative moving resistance between the components 3 and 4 is larger than the motion resistance of the slider, and it keeps the components 3 and 4 relatively static. When the mechanism is in configuration 2, a geometric constraint is added at kinematic joint E, so that the slider moves to the specified position and stops moving when the mechanism meets the geometric limit. When the mechanism is in configuration 3, the spring force at kinematic joint D makes the components 3 and 4 to be relatively static. Based on the principle of augmented Assur groups, the planar doublefolded metamorphic mechanism is divided into a fixedaxis rotating active part and an augmented Assur group RRRP (Fig.\u00a012b and c). 1 3 According to the geometric and physical parameters of the planar double-folded metamorphic mechanism in Table\u00a02, a three-dimensional model is established in SolidWorks, as shown in Fig.\u00a013. The initial location of component 4 is 227.73\u00a0mm, and the initial position of the component 1 is \u03c0\u00a0rad. Assuming that the active part 1 rotates at a constant speed of 6 r/min (motion period is 10\u00a0s), the dynamic simulation is carried out in SolidWorks virtual prototype environment, and the relationship between the driving torque and time of the planar double-folded metamorphic mechanism is obtained. When the planar double-folded metamorphic mechanism is in configurations 1 and 3, the mechanism under force constraint can be regarded as consisting of an active part and an augmented Assur group RRRP (as shown in Fig.\u00a012b and c). The dynamic analysis of the planar double-folded metamorphic mechanism in configurations 1 and 3 is shown in Fig.\u00a014. The dynamic equations in configurations 1 and 3 can be obtained by Eqs.\u00a0(1) and (4) as follows, When the mechanism is in configuration 2, the mechanism under geometric constraint can be regarded as consisting of an active part and an Assur group RRR, as shown in Fig.\u00a015. The dynamic analysis of the planar double-folded metamorphic mechanism in configuration 2 is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003253_10402004.2021.1958967-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003253_10402004.2021.1958967-Figure4-1.png",
        "caption": "Fig. 4: FEA model for mounting of the seal onto the shaft.",
        "texts": [
            " Simulation analysis methods Acc ep t d M an us cr ipt The presented design for a frictionless, elastomeric lip seal is examined in a finite-element analysis (FEA), in which the seal is mounted onto the shaft. Furthermore, an elastohydrodynamic lubrication (EHL) analysis is performed to study the seal\u2019s liftoff capability (i.e., sliding at near-zero friction on a thin air film when the shaft rotates). The assembly analysis is performed according to (12). Here, a coned shaft modelled as a rigid solid is quasi-statically inserted into an initially flat elastomeric ring (Fig. 4). A two-parameter Mooney\u2013Rivlin hyperelastic model ( 10 = 0.013 and 01 = 1.683) is used for modeling the fluoroelastomer (FPM) material. A 3D segment of the seal is used for the geometry model. Due to the high displacement during mounting, geometric nonlinearity effects are considered in the FEA. The surface roughness of the sealing material is excluded from the FEA model. An isothermal, isoviscous EHL analysis is performed to characterize the liftoff capability of the seal. The governing equations comprise the following:  A fluid-mechanics equation for the hydrodynamic pressure buildup in the lubricating film  A film thickness equation  A load balance equation The hydrodynamic pressure buildup is analyzed with the Reynolds partial differential equation (28) formulated in terms of the dimensionless density : \ud835\udc62s 2 \ud835\udf15 \u210e \ud835\udf15 \u2212 \ud835\udf15 \ud835\udf15 ( \u210e3 12\ud835\udf02 \ud835\udf15 \ud835\udf15 ) \u2212 \ud835\udf15 \ud835\udf15 ( \u210e3 12\ud835\udf02 \ud835\udf15 \ud835\udf15 ) = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001862_j.optlastec.2020.106714-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001862_j.optlastec.2020.106714-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram of laser equipment.",
        "texts": [
            " Flat tensile samples were cut from the brake camshaft of the trailer by electric spark machine, whose dimensions were gauge length of 24 mm, gauge thickness of 2 mm, and gauge width of 10 mm. The appearance of the specimen and bionic unit is given in Fig. 2. Seeking to obtain an ideal surface without pits as far as possible, the surface oil and wire-cutting processing marks of specimens were polished by using 80#, 160#, 320#, 600#, 800#, 1000# grades sandpapers. To be prepared for laser processing, and then samples were rinsed ultrasonically thoroughly with acetone and ethanol solution avoiding some accidental factors resulted from the uneven surface. Fig. 3 is a sketch of the Nd:YAG laser equipment with 1.064 \u03bcm wavelengths and a maximum power of 300 W, which was used to fabricate striation bionic units. The laser processing electric current was 160 A, pulse duration was 10.0 ms, laser frequency was 5 Hz, scanning speed was 10 mm/s and a circular spot size was 5 mm. Samples with 0\u25e6, 20\u25e6, 40\u25e6, 60\u25e6, and 90\u25e6 were also processed for investigating the effect of bionic unit\u2019s angular distribution on mechanical properties. At the same time, the distance of bionic unit is set as 4 mm based on the previous study, as illustrated in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000834_crc.2019.00024-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000834_crc.2019.00024-Figure7-1.png",
        "caption": "Fig. 7. Touch sensor on the forearm link",
        "texts": [
            " The robot has manipulators attached to the left and right joints of its upper half; each arm has 5 degrees of freedom (DOFs) and the hand has 1 DOF, for a total of 6 DOFs (Fig. 6). As configured for this study, the length of the upper arm link [from Joint 2 (shoulder) to Joint 4 (elbow)] was l2 and the length of the forearm link (from Joint 4 to Joint 6) was l4C. The manipulator joint angles were \u221290 deg \u2264 \u03c62 \u2264 +90 deg and 0 \u2264 \u03c63\u2264 +100 deg (Fig. 3). The right and left hands each had two fingers (Joint 6) to hold items (Fig. 6). Touch sensors were set on both forearm links, each of which were covered by bumpers (Fig. 7). These touch sensors were used when the robot front wheels climb a step (see Section III). The robot had a stopper mounted on its front of its body (Fig. 8 (a)). The stopper limited the passive rotational travel of the manipulators during wheelchair pushing (Fig. 8 (b)) and enabled the robot to imitate the operation of a human pushing an object (Fig. 9). The robot did not need to exert 76 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 06:18:12 UTC from IEEE Xplore",
            " The pushhandle mechanism included a rotary shaft that allowed passive rotation of the handle, and the robot hands grasped this shaft to connect the vehicles (Fig. 11 (a)). The wheelchair was equipped with a stopper composed of front and rear bars (Fig. 10) under the push-handle mechanism. During the climbing process, the arms are spread apart and the right and left manipulators are inserted between the stoppers (Fig. 11 (b)). As described above, the dual manipulators of the robot had touch sensors to detect contact with the stopper (Fig. 7). The robot pushed the wheelchair front bars using the forearm links of the manipulators to lift its front wheels (see Section III). The rear bars of the stopper prevented the robot from tipping over when the robot tilted backward as its mass position shifted behind the contact point between the robot middle wheels and the ground. Fig. 12 is a configuration diagram for the system controls. Here, the solid lines show the wired network, and the dotted lines show the wireless network. The robot has a camera, and a caregiver is able to watch a video feed, but in this study, the robot and wheelchair operated automatically to climb steps, and nobody needed to operate the system manually",
            " robot supports the wheelchair during this process to prevent the wheelchair from tipping over backward. <8> Both vehicles move forward, and the chair rear wheels climb the step. The level sensor on the wheelchair detects the completion of the wheelchair climbing onto the step (Fig. 14). Stage 3 <9> The system retracts the front and rear wheel mechanisms of the robot (Fig. 5). The arms are spread apart (Figs. 11 (a) and (b)), the touch sensors on the forearm links detect the contact with the stopper of the wheelchair (Fig. 7), and the manipulator forearm links are inserted into the stopper (between the front and rear bars). The chair stops, the robot moves forward, and the manipulator forearm links come into contact with the front bars of the wheelchair stopper (Fig. 10). <10> The robot continues to push the wheelchair using the forearm links, and the front wheels of the robot are lifted until the accelerometer system on the robot detects the robot inclination indicating that both vehicles are in a position where the robot\u2019s front wheels are on the step",
            " The robot prevented the wheelchair from tipping over backward, and it supported the ascent of the wheelchair rear wheels. After rear wheels of wheelchair reached the step, the level sensor on the chair detected the end of wheelchair\u2019s ascent. In stage 3, the wheelbase of the robot was changed to WB f = 220 mm (Fig. 3) to achieve the configuration needed for robot climbing. The mechanisms of the front and rear wheels were retracted (Fig. 5), The arms are spread apart, the touch sensors on the forearm links detected contact with the stopper of the wheelchair (Fig. 7), and the dual manipulators were set in the wheelchair stopper (Figs. 11 (a) and (b)). The robot detected the position for lifting the robot front wheels. The wheelchair maintained its position as the robot moved forward. The accelerometer on the robot detected the robot angle for step climbing, and the robot front wheels touched the step. The wheels were affected by the force from the step and the wheelbase, WB f , was shortened so that the front wheels could be placed on the step. In stage 4, both vehicles continued to moved forward, and the robot middle wheels contacted the step"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000266_ccac.2019.8920886-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000266_ccac.2019.8920886-Figure1-1.png",
        "caption": "Fig. 1. Ball & Beam model.",
        "texts": [],
        "surrounding_texts": [
            "Index Terms\u2014Nonlinear control, Passive system, Feedback linearization, Sliding mode control, Ball and Beam system.\nI. INTRODUCTION\nThe ball and beam system is a typical platform to implement and study the results of different modern control algorithms. The control structure of this system is also utilized for many different schemes in practical applications, e.g., it is used to demonstrate physical control application such as aircraft rollyaw regulation [1]. The ball and beam construction consists of two segments which are ball and beam body, system for motor control and has two degrees of freedom (DOF). It is hard to control the position of the ball because the ball moves continuously on the beam. In addition, because of the ball movement increase without limit for an angle of the beam, the system becomes open-loop unstable. [2]\nThe nonlinearity in the Ball and Beam has been studied in control theory since a considerable part of the real systems behaves according to a nonlinear dynamic. One of the problems to the stabilization of the system came from the unstable centrifugal force. Therefore, the concept of passivity as an input-output property allows a physical understanding about energy exchange in the systems and is one of the ideas in systems and control theory that has been used as a fundamental design tool for linear and nonlinear systems [3].\nThe development of the control has to include nonlinear feedback to stabilize the positions of the ball on the beam while tracking the desired reference signal. Researchers has used several control techniques such as: linear ProportionalIntegral-Derivative (PID) Control, linear quadratic regulator (LQR) [4], Adaptive control [5], Fuzzy PID control [2] [6], sliding mode control (SMC) [1] [7] [8], robust control [9], .\nThe sliding mode control has been used in different kind of systems due to provide good robustness properties, and it\napplied to both simple and complex structure. Creating the surface of control law can be applied the method of feedback linearization that transform the coordinates nonlinear system into a linear system whose states are the n \u2212 1 derivatives of the output variable that allows to the SMC neutralize the nonlinearities that cause the instability in the control law.\nThe Figure (1), shows the schematic of the ball and beam designed in specialized software to simulations.\nThis work proposes an analysis from the passivity with the purpose to demonstrate that the storage function can be related to the candidate Lyapunov function and the conditions to the stability. That allows design a sliding mode control based on feedback linearization to stabilize the position of the ball on a beam refusing the centrifugal term to reduce the model and simplify the severe nonlinearity.\nThis paper is organized as follow. The Section II, presents the dynamic model of ball and beam with passivity approach, and the sliding mode control is described too. The Section III shows the results of the simulations for the SMC and the comparison with a linear PID controller. Finally the Section IV concludes the paper.\n978-1-5386-6962-4/19/$31.00 \u00a92019 IEEE",
            "II. METHODOLOGY\nThe nonlinear equations that described the dynamic of the system can be written as :(\nJb r2 +m\n) p\u0308\u2212mp\u03b8\u03072 +mg sin \u03b8 = 0 (1)\n2mpp\u0307\u03b8\u0307 + (mp2 + J)\u03b8\u0308 +mgp cos \u03b8 = \u03c4 (2)\nWhere the system parameters of (1) and (2) are showing in Table I.\nSimplifying the model and refusing the centrifugal acceleration term p\u03b8\u03072 of (1). Let the states of the system as can be seen in (3):(\nJb r2 +m ) x\u03072 \u2212 : 0 mx1x 2 4 +mg sinx3 = 0\n(mx21 + J)x\u03074 + 2mx1x2x4 +mgx1 cosx3 = \u03c4\n(3)\nwhere the term : 0\nmx1x 2 4 is neglected according to [7].\nThe nonlinear state space of the ball and beam is showing in (4).\nx\u03071 x\u03072 x\u03073 x\u03074\n =  x2 \u2212mg sin(x3) Jb r2 +m\nx4 \u22122mx1x2x4\u2212mgx1 cos(x3)\nmx2 1+J\n+  0 0 0 1\nmx2 1+J\nu (4)\nConsider the nonlinear state space:\n\u03a3 = x\u0307 = f(x) + g(x)u\ny = h(x, u) (5)\nwhere f(x) and g(x) are functions that contains the nonlinear model.\nDefinition 1: Passive system [10]. The system \u03a3 is said to be passive if it is dissipative with respect to the supply rate w(t) = uT (t)y(t) and there exist a non negative real function H(x) called storage function, such that:\nH\u0307(x) \u2264 uT y (6)\nThe interpretation of the above definition is the energy increase (storage function) is no greater than the power inputoutput to it. The storage function is continuously differentiable, i.e., it is C1 and the passive system satisfies H(0) = 0.\nDemonstrating that the ball and beam is passive, first of all we can expressed the system equations into the lagrangian dynamic matrices form [3]:\nD(x)x\u0307+ C(x1, x2, x4)x+ g(x1, x3) =Mu (7)\nWhere:\nD(x) =  0 0 0 0 0 ( Jb r2 +m )\n0 0 0 0 0 0 0 0 0 (mx21 + J)\n ;\nC(x) = 2 3m  0 0 0 x1x2 0 0 0 0 0 x1x4 0 0\nx2x4 0 0 0\n ;\ng(x) = (\n\u2202V(x) \u2202x )T With V = mgx1 sinx3 being the potential energy.\nDefining the energy store function in (8):\nH = 1\n2 xTDx+ V(x) (8)\nApplying the derivative to (8), we obtain (9).\nH\u0307 = xTDx\u0307+ 1\n2 xT D\u0307x+ \u2202V \u2202x x (9)\nSolving (7) in function of the term Dx\u0307 and replacing in (9) we have (10):\nH\u0307 = xT [Mu\u2212 Cx\u2212 g(x)] + 1\n2 xT D\u0307x+ ( \u2202V \u2202x ) x (10)\nWhere \u2202V \u2202x = gT (x). Reorganizing (10), we have (11):\nH\u0307 = 1 2 xT [D\u0307 \u2212 2C]x\u2212 :xT g(x) + : gT (x)x + xTMu (11)\nWhere xT [D\u0307 \u2212 2C]x = 0 is the skew-symmetry property, therefore its quadratic matrix form is equal to zero and the term can be eliminated from (11).\nNow, the passive inequality is given by (12):\nH\u0307 \u2264 uMTx (12)\nWith MT = [1 0 0 0]. Therefore, the passive input-output map can be written as \u03a3 : u\u2192 x1.\nThe passivity inequality satisfies H(0) = 0, that can demonstrate the stability similar to the Lyapunov criteria, where the storage function is semidefinite negative and can be a candidate Lyapunov function being stable in open loop. The Ball and Beam is a passive system and to achieve the asymptotic stability is necessary a control loop."
        ]
    },
    {
        "image_filename": "designv11_63_0000466_tee.23094-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000466_tee.23094-Figure6-1.png",
        "caption": "Fig. 6. Flux line and distribution of current density. (N r = 10 000 rpm, f = 1 kHz, I max = 51 A). (a) Without magnetic composite material; (b) With magnetic composite material",
        "texts": [
            " 3 is sufficiently small compared with the reduction of copper loss. Losses can be reduced by attaching the magnetic composite material to the windings when considering the electric loss of the motor. Compared with when the magnetic composite material is not attached, the torque of each pasted pattern slightly increased. Therefore, the motor constant K m hardly changes. This is due to the improvement of the magnetic flux density in the slot teeth, which is caused by the leakage magnetic flux flowing through the magnetic composite material to the slot teeth. Figure 6 shows the current density distribution without magnetic composite material attached and for Pattern No. 3. In the case of Pattern No. 3, the number of magnetic fluxes that interlinked with the copper wire was small, and the bias of the current density was reduced. This is because magnetic composite material reduces the number of magnetic fluxes interlinked with the windings. Software JMAG-Designer (\u00d764) Ver.17.0 Analysis method Two-dimensional magnetic field analysis Solution FEM Mesh size 1) Copper: 1/10 or less of the skin depth 2) Magnetic layer: Automatic 3) Air: Automatic Analysis area Analysis in 10 times the analysis model Rotor Speed N r = 10 000 rpm Frequency Fundamental frequency: f = 1 kHz Current I = 51 Amax Material 1) Copper: (\u03c1 = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000993_s106879982001002x-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000993_s106879982001002x-Figure2-1.png",
        "caption": "Fig. 2. Sensor installation diagram.",
        "texts": [
            " The other end of the shaft rotates freely in the bearing. The coupling bearing is loaded with a constant radial force of 250 N applied to the nosepiece. The coupling and shafts rotate at an angular speed equal to the nominal speed, n = 2684 rpm. At the test bench, the angle of coupling axes bend is set at 10 5\u2032\u00b1 . During the experiment, the gear coupling is not loaded by torque. The radial acceleration sensors of the frame 1z and the bearing 2z are mounted on the adapter flange and on the outer race of the bearing through the bracket (Fig. 2). The radial directions of the sensor axes are the same. Thus, the sensors register the radial accelerations of parts separated by a rubber support. Such installation of two sensors allows further subtraction of the radial acceleration of the frame 1z from the radial acceleration of the bearing 2z in order to separate the dynamic processes in the transmission from disturbances that are caused in flight by forced vibrations of the fuselage and tail boom. To record the information, the equipment was used with four recording channels and a signal quantization frequency of 2000 Hz"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000792_ica-200622-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000792_ica-200622-Figure3-1.png",
        "caption": "Fig. 3. Layout of the cart equipped with an RGB-D sensor.",
        "texts": [
            " Figure 1 provides a general view of the proposed circular motorised rail system showing the initial configuration of the carts and the final configuration of the carts after using the intelligent trajectory planner and Table 1 shows a description of the different sym- bols that will be introduced and described along this research work. Figure 2 illustrates the general scheme of the platform, which consists of the following components: 1. A motorised cart system composed of two carts mounted on the circular rail and equipped with an RGB-D sensor. The aim of this system is to relocate the two motorised carts to monitor the patient\u2019s face and the physical rehabilitation exercise from two complementary viewpoints. 2. Two motorised carts equipped with an RGB-D device each (see Fig. 3), which move on the rail in a continuous manner to register the rehabilitation exercises and the facial expressions of the patient who freely moves inside the circle formed by the rail. 3. An intelligent trajectory generator that generates the coordinate trajectories of the two carts to locate them in the optimal configuration to monitor the patient. The intelligent trajectory generator comprises the references required to move the carts (denoted as x\u2217Q(t) and x\u2217S(t)) based on the instantaneous position of the patient in the circular rail (denoted as P(t)), the instantaneous position of the part of the body to be monitored in the circular rail (denoted as Pb(t)), the direction the person is looking at (denoted as principal vector, vp(t)), the direction to which the exercised part of body is pointing (denoted as secondary vector, vs(t)), and the instantaneous positions of the two carts (denoted as xQ(t) and xS(t))",
            " In order to perform the reference trajectories required to move the carts, it is necessary to establish the role (master or slave) of each cart, the direction (clockwise or counter-clockwise) that each cart must follow, and to generate the trajectories that avoid collisions between both carts. 4. A GPI control action for each cart that uses the reference trajectories constructed by the trajectory generator, generating the necessary control actions in order to achieve both the stabilisation of the system and the tracking of the references (i.e. xQ(t)\u2192 x\u2217Q(t) and xS(t)\u2192 x\u2217S(t)) and, consequently, to correctly monitor the patient (see Fig. 2). The dynamic model has been developed using as a basis the cart system shown in Fig. 3. The cart incorporates smooth support wheels, which are mobile supports that enable the cart to move, along with a rack and pinion that transmits its movement. The cart has been divided into three subsystems, which will make it possible to compute the balance of forces. These subsystems are the RGB-D sensor, the body of the cart and the sprocket. The calculations first consider the RGB-D sensor, after which they follow the transmission of forces to the motor that moves the sprocket. The dynamic model is obtained by attaining the function responsible for providing the torque generated by the motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002120_s00170-020-06571-5-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002120_s00170-020-06571-5-Figure7-1.png",
        "caption": "Fig. 7 Pinion-machining coordinate system",
        "texts": [
            " The normal curvature KP lineP, the geodesic torsion GP lineP along the vector tPline, and the normal curvature KP verticalP along the vector tPvertical at the pinion mean contact point, which are modified according to the contact attributes, are calculated as follows, KP lineP \u00bc KW lineP\u2212K WP lineP KP verticalP \u00bc KW verticalP\u2212K WP verticalP GP lineP \u00bc GW lineP\u2212G WP lineP 8< : \u00f041\u00de By this, the first-order parameters of each point on the contact path, and the first-order and modified second-order parameters at the mean contact point of the pinion concave flank are determined. The first-order parameters of the mean contact point of the pinion convex flank are determined, as well. Figure 7 shows the pinion-machining coordinate system, with oM \u2212 xMyMzM as the fixed machine-tool coordinate system, oc \u2212 xcyczc as the pinion cutter coordinate system, oPd \u2212 xPdyPdzPd as the pinion-fixed coordinate system without rotation with the pinion, and oP \u2212 xPyPzP as the pinion-moving coordinate system connected rigidly with the pinion. Point oP and point oPd coincide with the pinion crossing point, and the xP-axis and xPd-axis coincide with the pinion axis. A is the machine root angle, and (XOc, YOc, ZOc) is the position coordinate of pinion cutter center oc, expressed in oM \u2212 xMyMzM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure14-1.png",
        "caption": "Fig. 14. Temperature distribution with axial water jacket cooling for J= 23.3 A(rms)/mm2",
        "texts": [
            " The channels are removed from the winding supports and they are replaced with solid winding supports made of D5506 material. Similar to the proposed cooling method, the water jacket housing is made of aluminium and has one axial channel per slot for a total of 12 channels. Since only one channel carries the coolant, the entire flow volume is through this channel, 1087 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 19,2021 at 14:42:00 UTC from IEEE Xplore. Restrictions apply. thus keeping the total flow rate the same as discussed in Section III. As presented in Fig. 14, the maximum temperature is well above the permissible limits. Thus, the maximum allowable slot current density using axial water jacket cooling for the same temperature rise of 80 \u25e6C, as seen in Fig. 15, is 15.6 A/mm2. This current density corresponds to 50% lower compared to that of the WELC method. This paper proposes a low thermal resistance winding embedded liquid cooling (WELC) concept for slotless motors leveraging the space within the non-magnetic winding support for efficient heat extraction out of the windings, and thereby achieving higher current densities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002929_j.addma.2021.102072-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002929_j.addma.2021.102072-Figure1-1.png",
        "caption": "Fig. 1. (a) The design of the demolding test tool and schematic drawing of applied forces between the tool and part during the demolding process. (b) The demolding test tool manufacturing process; computer aided design (CAD) model, slicing image, after additively manufactured, and after post-machining process of the tool.",
        "texts": [
            " Further, the tool surface roughness and surface defects were analyzed in multiple composite part manufacturing cycles to investigate the sustainability of the coating on the additively manufactured composite tool. G. Kim et al. Additive Manufacturing 46 (2021) 102072 An experimental procedure was developed to investigate the effects of the coating on the demolding characteristics of printed composite tools, namely the demolding force and surface roughness. To quantify the effects of the coating on the demolding force, a tool geometry was designed to promote a demolding force resulting from the friction between the part and the tool. Fig. 1(a) shows the cross-section of the tool for a cylindrical shape designed for the experiments used for characterizing the demolding force. A composite prepreg material was laid at the top and sides of the tool. When the composite material was cured in an elevated temperature and cooled down, the composite material shrank and generated normal force against the tool. The composite material with higher CTE shrank more and generated higher normal force. During the demolding process, the sliding composite part on the tool applied a frictional shear stress between the tool and the part",
            " The nozzle temperature was set to 300 \u25e6C for the printing and the printing speed was 5500 mm/s. The total printing time of the demolding test tool was 8 min. The printed tool was heat treated at 130 \u25e6C, which was about 20 \u25e6C above the glass transition temperature (Tg) of the printed material for 2 h for thermal stress relaxation [29,30]. The tool was post-machined using a CNC milling machine with 500 SFM and 0.005 IPT [19]. The tool was cleaned with isopropyl alcohol, dried, and inspected. An ejector plug was manufactured from aluminum 6061-T6. Fig. 1(b) shows the demolding test tool manufacturing process. For the coating material, a commercially available thermoset liquid coating with ceramic particles, Cerakote E-series, was used in this work. The primary motivation for adopting such a coating material was its thermal characteristics and simple application. First of all, the application temperature of the coating material was relatively low, 82 \u25e6C [19, 31]. This was important to prevent permanent deformation of the tool that could occur while curing the coating at temperature (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002239_s11771-021-4591-3-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002239_s11771-021-4591-3-Figure2-1.png",
        "caption": "Figure 2 Geometric model of planetary gear with inner gear",
        "texts": [
            " Meanwhile, based on the numerical solution of the EHL, a model for calculating oil film stiffness was established to characterize the dynamic properties of the oil film more clearly. The variation of the oil film stiffness under different rotation speeds and modification coefficients was investigated. As shown in Figure 1, a planetary gear train consisted of sun gear, planetary gear, inner gear, and planetary frame. The geometric model of the internal meshing gears in the planetary gear train is shown in Figure 2. N2 and N3 represent the theoretical engage-in point and recess point, respectively. Actually, A is the true engage-in point and D is the true recess point. rbp and rbr denote the radii of the base circles (p represents the planetary gear and r denotes the internal gear here and below) and their J. Cent. South Univ. (2021) 28: 126\u2212139 128 rotation speeds are np and nr, respectively. P is the pitch point, and K denotes any possible meshing point during the transmission process. For high-speed and heavy-duty gear systems, the damping ratio and inertial forces have significant effects on the transmission characteristics, and thus, dynamic analysis is necessary",
            " The dynamic load for the meshing teeth between the planetary gear and the inner gear is * d rp rp rp rpF k c    (6) The fluctuating velocity induced by the vibration of the inner gear is r br ru r   (7) The fluctuating velocity induced by the vibration of the planetary gear is J. Cent. South Univ. (2021) 28: 126\u2212139 129 p bp p c c rpcosu r r     (8) The dynamic component of the entrainment velocity due to the vibration can be written as r p( ) / 2u u u  (9) 2.3.1 Basic contact parameters for planetary and inner gears As depicted in Figure 2, according to the properties of the involute curve, the instantaneous radii of curvature at any possible meshing point K are as follows: p 2 2 bp rp( ) tan ( )R K N K N P PK r s t     , r 3 3 br rp( ) tan ( )R K N K N P PK r s t     (10) where s denotes the distance from the meshing point K to the pitch point P. The value of \u03b1rp can be obtained by the following equation: 3 2 3 2 rp(inv inv ) 2 tan z z x x x         (11) where z3 and z2 denote the tooth number of the inner gear and the planetary gear, respectively, and the modification coefficients are denoted as x3 and x2, respectively",
            " Figure 7 shows the variations of the comprehensive radius of curvature, entrainment velocity, and slide-roll ratio. The modification coefficient had significant influence on the tooth profile. Among the different transmission types, both the comprehensive radius of curvature and the entrainment velocity were the largest when the positive drive was used. For the negative drive, the comprehensive radius of curvature was the smallest, leading to the smallest entrainment velocity. The inner gear was active. As shown in Figure 2, A was the engage-in point. Based on the properties of the involute curve, the comprehensive radius of curvature decreased consistently in the transmission process. Affected by the comprehensive radius of curvature, the entrainment velocity continued to decrease. J. Cent. South Univ. (2021) 28: 126\u2212139 132 Figure 7 Transmission characteristics of gear system: (a) Equivalent radius of curvature; (b) Entrainment velocity and slide-roll ratio Regardless of the type of transmission, over the whole process, the slide-roll ratio increased constantly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure8-1.png",
        "caption": "Fig. 8 Expert based design optimisation (EBDOP-1)",
        "texts": [
            "5-mm-thick upper panel, a 1.5-mm-thick lower panel and 2.5-mm-thick towers. These sub-components of the cross member are fastened to each other by electric resistance welding and arc welding. The baseline cross-member design is shown in Fig.\u00a07. One alternative to the first proposal design involves reducing only the depth by 10\u00a0mm compared to baseline. This change causes a mere 0.2\u00a0kg weight reduction per part. This proposal design is referred to as SBDOP in this paper. The EBDOP-1 design is shown in Fig.\u00a08. In a second alternative design, some cut-outs are made on the upper and lower panels, but the panels\u2019 thicknesses remain the same. This proposal design is referred to as EBDOP throughout this paper. The EBDOP-2 design is shown in Fig.\u00a09. During the optimisation, 70 iterations have been carried out to obtain the optimised design. The number of iterations used is related to the convergences of the problem, such as changes in the object function or the minimisation of mass. The only gauge change was a reduction of the lower panel thickness from 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002239_s11771-021-4591-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002239_s11771-021-4591-3-Figure1-1.png",
        "caption": "Figure 1 Structure of planetary gear train",
        "texts": [
            " In this work, a dynamic model of the planetary gear train was first established, and then the vibrations of the gear system were coupled with the lubrication, and the dynamic behaviors and lubrication performances considering different transmission types and modification coefficients were examined. Meanwhile, based on the numerical solution of the EHL, a model for calculating oil film stiffness was established to characterize the dynamic properties of the oil film more clearly. The variation of the oil film stiffness under different rotation speeds and modification coefficients was investigated. As shown in Figure 1, a planetary gear train consisted of sun gear, planetary gear, inner gear, and planetary frame. The geometric model of the internal meshing gears in the planetary gear train is shown in Figure 2. N2 and N3 represent the theoretical engage-in point and recess point, respectively. Actually, A is the true engage-in point and D is the true recess point. rbp and rbr denote the radii of the base circles (p represents the planetary gear and r denotes the internal gear here and below) and their J. Cent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002367_iros45743.2020.9341539-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002367_iros45743.2020.9341539-Figure3-1.png",
        "caption": "Fig. 3. A comparison of holomonic and non-holomonic robots: (a) locomotions, (b) safety envelopes, (c) collision areas of VO and ORCA, and (d) motion planning of ORCA.",
        "texts": [
            " Second, based on the range measurements and the exchanged information on relative velocities, the relative poses of the other robots in the coordinate system of robot i are estimated and updated. Finally, the optimal velocity that guarantees the collision-free motion of the robot is chosen by using NH-ORCA, given the input of relative poses of other robots with uncertainty models. NH-ORCA is an extension of ORCA that guarantees the collision free motion of multiple non-holonomic robots. The locomotions of holonomic and non-holonomic robots in the velocity of vH is shown in Fig. 3(a), which has a tracking error \u03betrack. The NH-ORCA includes three main steps. First, the tracking error is added with the robot radius to construct the safety envelopes and collision areas, as shown in Fig. 3 (b-c). Secondly, the set of allowed holonomic velocities SAHV is computed using the velocity vH to limit the tracking error bound. It is represented by a convex polygon PAHV . Finally, within the set of safe velocities computed by PAHV and ORCA, the optimal velocity closest to the desired velocity is chosen and mapped into the corresponding nonholonomic control inputs, as shown in Fig. 3 (d). For a group of non-holonomic robots navigating in the 2D environment, the desired velocity is the velocity that guides robots to the target position directly without the consideration of obstacles within each time window. The main problems of the range-only multi-robot collision avoidance include (1) how to estimate the robot poses with range measurements and exchanged information on velocities in a distributed way; (2) how to find a safe velocity for each robot to achieve a collision-free and time-efficient motion within a time horizon \u03c4 given the robot poses with uncertainties"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure1-1.png",
        "caption": "Fig. 1. Configuration of the conventional HTS-FS machine.",
        "texts": [
            " More importantly, the use of the single ring-shaped HTS-excitation coil can reduce the waste of HTS wires, and the radial partitioned stator scheme can relieve the influence of armature field on the HTS-excitation coil. In Section II, the configuration and operation principle of the proposed machine will be discussed. Section III will be devoted to the machine design, mainly focusing on the design of HTS-excitation coil and refrigeration system. And then, by using the three-dimensional finite element analysis (3D-FEA), the electromagnetic performances of the proposed machine will be presented and analyzed in Section IV. Finally, some conclusions are drawn in Section V. As shown in Fig. 1, in the conventional HTS-FS machine, the armature windings and HTS-excitation coils are placed on the same stator, namely, the HTS-excitation coil spans two adjacent 1051-8223 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 11:27:00 UTC from IEEE Xplore. Restrictions apply. stator teeth",
            " As the rotor moves further 180 degrees to 270 electrical degrees, the flux linkage of coil A1 will be at its negative peak as shown in Fig. 3(b). It indicates that in the proposed RPS-HTSFS machine, the total flux linkage of the phase A by connecting the armature coils in series can be bipolar. Thus, as the rotational movement passes one by one pole pitch, the electromotive force (EMF) of the armature winging changes periodically as well. For the HTS-excitation coil, there are two feasible winding methods, namely, ring-shaped and racetrack-shaped structures as depicted in Fig. 4. In the conventional HTS-FS machine as shown in Fig. 1, several racetrack-shaped HTS-excitation coils are employed, which brings longer end winding. On the other hand, in the winding process of HTS-excitation coil, the coil needs to be applied winding tension and then the bending stress is produced. When the winding tension is over large, the inner layer of the HTS-excitation coil will be excessive bending, causing the degradation of mechanical properties. Therefore, it is of great necessity for the HTS-excitation coil to consider the proper bend radius and try to reduce its bending stress"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001505_icecce49384.2020.9179221-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001505_icecce49384.2020.9179221-Figure1-1.png",
        "caption": "Fig. 1. Nomenclature of Four-Bar mechanism",
        "texts": [
            " In another study, the dimensional synthesis of body guidance and function generation, using the graphical and analytical methods was presented. The output results show that when two positions of the coupler are specified, there are infinite numbers of design solutions. If three positions of the coupler are specified and the positions of the moving joints are also given, then there is only one solution [21]. II. DESCRIPTION OF FOUR-BAR SYSTEM A four-bar mechanism is a simple mechanism that consists of four rigid links which are linked in the form of quadrilateral by four revolute joints. As shown in Figure 1, a four-bar mechanism with links lengths r1, r2, r3, and r4 and angles \u03f41, \u03f42, \u03f43, and \u03f44 is taken for further analysis. \u03f41 is zero in this case. The masses of the links are m1, m2, m3, and m4. The center of mass of link 2 is at a distance of rc2 from center O. The center of mass of link 3 is at a distance of rc3 from end A. The center of mass of link 4 is at a distance of rc4 from O\u2019. Link 1 is called ground link as it is fixed with no motion. Link 2 is the input link on which we can apply forces and torques",
            " Link 3 is called the coupler link and it has a complex motion that is combination of translation and rotation. Link 4 is normally called output link and it has only rotational motion. The mechanism of four-bar can be described as \u201cGrashof\u2019s mechanism\u201d if following condition is satisfied + < + (1) Where, s = length of the shortest link; l = length of the longest link; p = length of one other link; q = length of second link. Grounding the various links of this mechanism one by one will result in four inversions for the 4-bar mechanism given in Figure 1. By grounding the link adjacent to the shortest, we will get a Crank-Rocker mechanism. In a Crank-Rocker mechanism, the shortest link will have a complete 360 degrees rotation, and the link opposite to the shortest will have oscillatory motion. By grounding the shortest link, we will get a Double-Crank mechanism in which both the input and output links will have a complete 360 degrees rotation. By grounding the link opposite to the shortest link, we will get a DoubleRocker mechanism. In a Double-Rocker mechanism, both the input and output links will have the oscillatory motions",
            " Downloaded on August 31,2020 at 13:11:46 UTC from IEEE Xplore. Restrictions apply. A. Kinematics Kinematics is the branch of mechanics that is concerned with the motion of objects without reference to the forces which cause the motion. The motion of the four-bar mechanism can be described by three fundamental parameters i.e. position, velocity, and acceleration, as explained in the succeeding passages. A-I. Position analysis for four-bar mechanism For the position analysis of a four-bar mechanism, consider Figure 1 again. There are various techniques that can be used for the position analysis of a four-bar mechanism i.e., Freudenstein method, graphical method, and method based on complex algebra. For current study Freudenstein technique was utilized for this purpose. From Figure 1 we can write the following vector loop closure equation: = + = + (4) The above position vectors can be written in terms of the link lengths and angles as: ( + ) + ( + )= + ( + ) (5) From the previous vector equation, two scalar equations may be obtained by comparing the \u201ci\u201d and \u201cj\u201d components on both the sides: ( ) = + ( ) \u2212 ( ) (6) ( ) = ( ) \u2212 ( ) (7) After adding and squaring we get: + M + N = 0 (8) The coefficients L, M, and N are then estimated from: L = 2 \u2212 2 ( ) (9) M = \u22122 ( ) (10) N= + \u2212 + \u2212 2 ( ) (11) Equation 8 is known as Freudenstein equation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002104_s00202-020-01155-8-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002104_s00202-020-01155-8-Figure1-1.png",
        "caption": "Fig. 1. 3D diagram of the analysed machine",
        "texts": [
            " Moreover, the efficacy of using ferrite PMs against demagnetization risks could be boosted by adopting interior permanent magnet (IPM) topology, as reported in [9]. More so, it is established in [11] that there is an inverse proportionality between the used quantity of PM material in a machine and its corresponding magnet remanence value, exclusive of the ferrite magnet type which has a direct relationship; however, in all cases, the amount of the induced electromotive force as well as its resultant electromagnetic torque is a function of the magnetic remanence in a given permanent magnet machine. A 3D-diagram of the analysed machine is shown in Fig.\u00a01. Further, the studies in [12] reveal that the output electromagnetic power of a PM machine is dependent on the quantity of its magnetic remanence with a synchronized lesser mass of the magnetic material, i.e. the ideal choice for larger output power is a greater magnetic remanence and lesser active mass of magnet. Furthermore, the ferrite magnet is ideal in applications where low PM eddy current losses are required, since it gives the lowest amount of magnet eddy current loss as a result of its intrinsic resistivity worth [12]",
            "\u00a02 that the effect of magnetic material in terms of coercivity, remanence of the magnet, and the resultant energy product (BH)max properties are more noticeable on both neodymium-based and samariumequipped machines under open-circuit and normal operating conditions, due to their higher value of both remanence and energy products which characterize both the intensity of the linking flux and the energy concentration of the material, respectively. For ease of writing, the employed neodymium\u2013iron\u2013boron material in this study would be abbreviated as neodymium, henceforth. Both 2D and 3D FEA time-stepping methods are employed in predicting the results; the analysed FEA model is depicted in Fig.\u00a01. It is worth noting that the used MAXWELL finite element software package automatically generates the basic electromagnetic results, such as flux linkages, electromotive force (EMF), and torque. Then, other machine parameters are obtained from post-processing of the generated results, on appropriate application of the simulation conditions, such as electric loading and no-load settings. Further, the fast Fourier transforms (FFT) of the generated FEA waveforms and results are computed using MATLAB program"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001568_acs.langmuir.0c02021-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001568_acs.langmuir.0c02021-Figure2-1.png",
        "caption": "Figure 2. (a) PS particle polarization model. (b) 3D polarization model of the PS particle. (c) Electric field gradient nephogram around the electrode. (d) Partial enlarged drawing of the electric field gradient nephogram. (e) Simulated potential curve of the interdigital electrode. (f) Change of Re[f CM] with frequency for PS particles in deionized water and 30 wt % acrylamide solution.",
        "texts": [
            " The picture of the fabricated gel is shown in Figure S1b. \u25a0 RESULTS AND DISCUSSION Assembly Mechanism under an AC Electric Field. Induced by an external electric field, the particle will be polarized, and the positive and negative charge centers will no longer coincide, forming an electric dipole.36 The direction of the equivalent electric dipole moment generated is along the direction of the electric field.37,38 COMSOL Multiphysics was used to simulate and analyze the polarization effect of a single particle. As shown in Figure 2a and Figure 2b, the PS particle with a diameter of 5 \u03bcm was placed in a nonuniform electric field. It can be found that the PS particle generates a polarization charge, and the polarization degrees on the left and right sides of the particle are not the same. Besides, the side with higher electric field intensity is more polarized, since red represents higher polarization intensity and blue represents lower polarization intensity. In the experiments, the electric field is provided by an interdigital electrode.39 The electric field intensity is also simulated by COMSOL, and the specific parameters involved are displayed in the Supporting Information. From the electric field gradient nephogram around the electrode (Figure 2c, Figure 2d), we can clearly see that the field intensity reaches its maximum value at the electrode edge. For exploring the distribution of the electric potential, the intercept line in the longitudinal section is taken every 10 \u03bcm above the electrode. The distance along the x-axis direction of the intercept line is taken as the abscissa, and the value of the electric potential is taken as the ordinate to draw the curve of potential changing with distance. As shown in Figure 2e, it can be seen that whenH = 0, the potential at the electrode surface can reach the maximum value of 10 V. As the height increases, the potential decreases. Moreover, the attenuation of the electric potential is not uniform, and the attenuation speed decreases with the increase of height. In other words, the closer to the electrode, the more obvious the potential attenuation. The force of an electric field on particles is the called dielectrophoresis force,40 and it is defined as follows: \u03c0\u03b5= [ ]\u2207F r f E2 ReDEP m 3 CM 2 (1) where \u03b5m is the absolute permittivity of the medium; r is the radius of spherical particle; E is the field intensity; \u2207E2 is the gradient of the electric field square; and Re[f CM] is the real part of f CM",
            " The effective polarizability for the spherical particle in medium is \u03b5=K f3 m CM (4) In combination with eq 2, 3, and 4, it can be concluded that \u03b5 \u03c0 \u03b5 \u03b5 \u03c3 \u03c3 \u03c0 \u03b5 \u03b5 \u03c3 \u03c3 = \u2212 \u2212 \u2212 + \u2212 + K f j f j 3 2 ( ) ( ) 2 ( 2 ) ( 2 )m p m p m p m p m (5) https://dx.doi.org/10.1021/acs.langmuir.0c02021 Langmuir 2020, 36, 11546\u221211555 11548 The direction of the DEP force depends on permittivities and conductivities of the medium and particles. At low frequency, the direction is mainly determined by conductivities. On the contrary, at high frequency, the direction is mainly determined by the permittivities. If Re[f CM] > 0, the particles will be subjected to a positive DEP force. If Re[f CM] < 0, the particles will be subjected to a negative DEP force. Figure 2f shows the change of Re[f CM] with frequency for PS particles in deionized water and 30 wt % acrylamide solution. Here, the relative permittivity and conductivity of PS particles are \u03b5p = 2.5, \u03c3p = 1\u00d7 10\u22123 S/m; those of deionized water are \u03b51 = 78, \u03c31 = 2 \u00d7 10\u22124 S/ m; those of acrylamide are \u03b52 = 60, \u03c32 = 5\u00d7 10\u22123 S/m. It could be found that when the frequency is more than 100 kHz,21 the force changes from positive DEP force to negative DEP force. Here, our experimental results are based on the negative dielectrophoresis behavior of the particles",
            " Induced by the combined action of the electric field and the electric double-layer, the vortex as shown in Figure 3a will be generated on the electrode surface, and the particles will move with the vortex. This phenomenon is alternating current electroosmosis (ACEO),41 which occurs on the surface of the electrode and is symmetrical along the center of the electrode. As a consequence of the symmetry, the flow rolls on the center of electrode are identical and cancel each other. Therefore, when the particle moves to the center of the electrode, it is easily captured and no longer moves. Furthermore, as can be seen from Figure 2d, the electric field lines converge or diverge in quantities at the electrode edge, and the electric field changes violently. The field intensity gradient is high, and the electroosmosis velocity reaches the maximum at the electrode edge. Therefore, the particles near the electrode surface will accelerate to converge to the center of the electrode and be captured due to the action of ACEO, while the particles far away from the electrode surface form a linear chain along the electric field between the electrodes, as shown in Figure 3b",
            " The particles in the lower layer seem to be sparse because their arrangement direction is perpendicular to the upper layer\u2019s. In the experiment, electrodes of different sizes were used to characterize the effect of voltage on equilibrium height measured by SEM. As shown in Figure 4e, it can be found that the height increases with voltage, and the increasing speed is related to the electrode width. The larger the electrode width, the faster the height increases. In addition, the electric field frequency also affects the equilibrium height, and the influence mechanism mainly includes two aspects. First, as shown in Figure 2f, when the frequency is greater than the critical frequency of p-DEP and n-DEP, Re[f CM] increases with the increase of frequency and tends to be stable when the frequency reaches about 1 MHz. Consequently, the DEP force increases first and then remains unchanged theoretically. Therefore, the equilibrium height should also increase first and then remain stable. However, under experimental conditions, the internal resistance and capacitance effect of the experimental device weaken the actual voltage loaded on the electrode, and with the increase of frequency, the effect is stronger and stronger"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001041_iet-epa.2020.0237-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001041_iet-epa.2020.0237-Figure5-1.png",
        "caption": "Fig. 5 Half of pole of the rotor",
        "texts": [
            " (10) where lav is the average conductor length of half a turn (m), Ns is the number of the stator slots, kcu is the thermal conductivity of the copper (W/m-K), Scu is the total copper conductor cross-section area (m2) Conduction resistance between the end windings and the motor housing (R7): The distance between the outer surface of the end winding and the internal part of the machine housing is low [20]; accordingly, the airspeed in this region is very low, and the heat exchange from the end winding to the internal part of the housing is assumed as conduction and is calculated as R7 = ln ros/(ros \u2212 \u03b3Hy) 2\u03c0kair Lec \u2212 Ls , (11) where Hyis the stator yoke height (m), \u03b3 is the reduction coefficient to implement the distance between the outer surface of the end winding and the internal part of the housing and Lec is the motor housing length (m). The value of \u03b3 for ordinary induction machines with single layer windings is defined in the range of 0.4\u20130.7 [17]. However, this value for the machine by the double layer windings decreases to the range of 0.25\u20130.4 [3]. Conduction resistances of the rotor (R11 and R12): Fig. 5 illustrates the cross-section of a one eighth of the rotor. The rotor structure is formed from the steel laminations, as the magnetic flux paths and air gaps in the laminations as magnetic flux barriers. As the heat inside the flux barriers is transferred by the natural convection and according to the high thermal conductivity of the steel laminations, consequently, the heat transfer by the natural cooling from barriers is eliminated and assumed whole of the heat is transferred by the conduction mechanism through the steel laminations. Fig. 6 illustrates the LP circuit of the rotor by implementing the above simplifying hypothesis. According to Fig. 5, the total value of the rotor resistance equals Rr, which consists of the seriesparallel combinations of thermal resistances. To implement the rotor losses to the LPTN, Rr is divided into two equal resistances R11 and R12 illustrated in Fig. 5 and the LPTN of the SynRM. The value of the thermal resistances presented in Fig. 6 is calculated by (1) and (3). The values of R1r and R11r to R15r are calculated by (3), and the rest are evaluated by (1). Tables 4 and 5 provide a better overview of the computation of the conduction resistances. Finally, the value of Rr is calculated by the series-parallel resistances law in the electrical circuit. Further, R11 and R12 are defined as IET Electr. Power Appl., 2020, Vol. 14 Iss. 10, pp. 1828-1836 \u00a9 The Institution of Engineering and Technology 2020 1831 Authorized licensed use limited to: Newcastle University"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000771_iccia49288.2019.9030988-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000771_iccia49288.2019.9030988-Figure8-1.png",
        "caption": "FIGURE 8. Power test bench of the CPI gyrotron. The gyrotron output power is the sum of the power of waveguide dummy load, aluminum tank dummy load, bellows, waveguides, waveguide switch, MOU, miter bends, and output window.",
        "texts": [
            " And the coupling of the power monitor miter bend for Gycom gyrotron is not clear, but it will not be the difficulty for us to realize the real time power monitoring as shown below. VOLUME 4, 2016 2169-3536 2016 IEEE. Translations and content mining are permitted for academic research only. Personal use is also permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. 5809 A. INCIDENT WAVE POWER MONITORING The 3D pictures of the power test benches for Gycom gyrotron and CPI gyrotron are shown in Fig.6 and Fig.8. One power monitor is situated between the MOU and the dummy load for CPI gyrotron. This power monitor miter bend is situated before the polarisers and it picks up only 5810 VOLUME 4, 2016 one polarisation. The output wave from the MOU is H-plane linearly polarised. As shown in Fig.6, two powermonitors are used for Gycom gyrotron. One is situated between the MOU and the dummy load, the other one which is similar to the one used for CPI gyrotron is situated between the first power monitor miter bend and tokamak"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002338_012093-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002338_012093-Figure6-1.png",
        "caption": "Figure 6. Element with force components in nodes and vector of equivalent force \ud835\udc46\ud835\udc46\ud835\udc3f\ud835\udc3f and vector of displacement \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f.",
        "texts": [],
        "surrounding_texts": [
            "For building an entire system of structure is required a matrix transformation. In this case, all sizes into the local system coordinates must be transformed into sizes of the global system coordinates. Switching from local coordinates to global coordinates for a node, also knowing the rotation angle of axis, we have: \ufffd \ud835\udc4b\ud835\udc4b \ud835\udc4c\ud835\udc4c \ud835\udc4d\ud835\udc4d \ufffd = \ufffd \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 \ufffd \u2219 \ufffd \ud835\udc51\ud835\udc51 \ud835\udc66\ud835\udc66 \ud835\udc67\ud835\udc67 \ufffd \ud835\udc4e\ud835\udc4e\ud835\udc60\ud835\udc60\ud835\udc51\ud835\udc51 \ufffd \ud835\udc51\ud835\udc51 \ud835\udc66\ud835\udc66 \ud835\udc67\ud835\udc67 \ufffd = \ufffd \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 \ufffd \u2219 \ufffd \ud835\udc4b\ud835\udc4b \ud835\udc4c\ud835\udc4c \ud835\udc4d\ud835\udc4d \ufffd (7) \ud835\udc45\ud835\udc45\ud835\udc41\ud835\udc41 = \ufffd \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 \ufffd (8) Switching from local coordinates into global for an element we have: \u239d \u239c \u239c \u239b \ud835\udc4d\ud835\udc4d1 \ud835\udc4d\ud835\udc4d2 \ud835\udc4d\ud835\udc4d3 \ud835\udc4d\ud835\udc4d4 \ud835\udc4d\ud835\udc4d5 \ud835\udc4d\ud835\udc4d6\u23a0 \u239f \u239f \u239e = \u239d \u239c\u239c \u239b \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 \u23a0 \u239f\u239f \u239e \u2219 \u239d \u239c\u239c \u239b \ud835\udc67\ud835\udc671 \ud835\udc67\ud835\udc672 \ud835\udc67\ud835\udc673 \ud835\udc67\ud835\udc674 \ud835\udc67\ud835\udc675 \ud835\udc67\ud835\udc676\u23a0 \u239f\u239f \u239e (9) \ud835\udc45\ud835\udc45 = \u239d \u239c\u239c \u239b \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \u2212 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 \ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60 \ud835\udefc\ud835\udefc 0 0 0 1 \u23a0 \u239f\u239f \u239e (10) \ud835\udc4d\ud835\udc4d\ud835\udc3a\ud835\udc3a\ud835\udc5d\ud835\udc5d\ud835\udc5f\ud835\udc5f\ud835\udc3a\ud835\udc3a\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc67\ud835\udc67\ud835\udc5d\ud835\udc5d\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d \ud835\udc4e\ud835\udc4e\ud835\udc60\ud835\udc60\ud835\udc51\ud835\udc51 \ud835\udc67\ud835\udc67\ud835\udc5d\ud835\udc5d\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38\ud835\udc47\ud835\udc47 \u2219 \ud835\udc4d\ud835\udc4d\ud835\udc3a\ud835\udc3a\ud835\udc5d\ud835\udc5d\ud835\udc5f\ud835\udc5f\ud835\udc3a\ud835\udc3a\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d (11) 10th EASN 2020 IOP Conf. Series: Materials Science and Engineering 1024 (2021) 012093 IOP Publishing doi:10.1088/1757-899X/1024/1/012093 Where \ud835\udc4d\ud835\udc4d and \ud835\udc67\ud835\udc67 are the vectors in global coordinates and local coordinates respectively. \ud835\udc46\ud835\udc46\ud835\udc3a\ud835\udc3a = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f; \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38\ud835\udc47\ud835\udc47 \u2219 \ud835\udc46\ud835\udc46\ud835\udc3a\ud835\udc3a \ud835\udc4e\ud835\udc4e\ud835\udc60\ud835\udc60\ud835\udc51\ud835\udc51 \ud835\udc37\ud835\udc37\ud835\udc3a\ud835\udc3a = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f; \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38\ud835\udc47\ud835\udc47 \u2219 \ud835\udc37\ud835\udc37\ud835\udc3a\ud835\udc3a (12) \ud835\udc46\ud835\udc46\ud835\udc3a\ud835\udc3a = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38\ud835\udc47\ud835\udc47 \u2219 \ud835\udc37\ud835\udc37\ud835\udc3a\ud835\udc3a (13) \ud835\udc46\ud835\udc46\ud835\udc3a\ud835\udc3a = \ud835\udc58\ud835\udc58\ud835\udc3a\ud835\udc3a \u2219 \ud835\udc37\ud835\udc37\ud835\udc3a\ud835\udc3a \u27f9 \ud835\udc58\ud835\udc58\ud835\udc3a\ud835\udc3a = \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38 \u2219 \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc45\ud835\udc45\ud835\udc38\ud835\udc38\ud835\udc47\ud835\udc47 (14)"
        ]
    },
    {
        "image_filename": "designv11_63_0002367_iros45743.2020.9341539-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002367_iros45743.2020.9341539-Figure1-1.png",
        "caption": "Fig. 1. An illustration of distributed multi-robot collision avoidance with range measurements only.",
        "texts": [],
        "surrounding_texts": [
            "orientations. The information on relative poses among robots is the precondition of all collision avoidance approaches. How to estimate the relative poses between robots with high accuracy from low data-throughout measurements is still an open question. 3) Robust collision avoidance with uncertainties in pose estimation. The uncertainties in relative pose estimation increase the risk of robot collision. It is vital to develop a robust collision avoidance scheme. Various approaches have been developed to solve these challenges such as reinforcement learning (RL) based approaches [2]\u2013[7], velocity obstacle (VO) based approaches [8]\u2013[11] and energy-optimal based approaches [12], [13]. Compared with RL and energy-optimal based approaches, VO based methods have lower computational complexities and are convenient for real-time implementation. Most centralized approaches assume that all the robots have a global coordinate system, or each robot has the perfect sensing capability of acquiring the accurate pose information on others [14], [15], an extra tracking system is often required in the real world experiments [16]\u2013[18], incurring high computational costs and limits upon applications in outdoor environments. For distributed approaches, each robot needs to estimate its own pose first using on-board sensors; then, the estimated poses are shared through inter-robot communications [19]\u2013[21]. Usually, expensive high-resolution sensors are unsuitable for applications of large-scale robot groups. Outdoor applications of those sensors highly sensitive to the changing environmental conditions are also limited. In this paper, a distributed multi-robot collision avoidance approach is proposed with range-only measurements using 978-1-7281-6212-6/20/$31.00 \u00a92020 IEEE 8020 20 20 IE EE /R SJ In te rn at io na l C on fe re nc e on In te lli ge nt R ob ot s a nd S ys te m s ( IR O S) | 97 8- 1- 72 81 -6 21 2- 6/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ IR O S4 57 43 Authorized licensed use limited to: Carleton University. Downloaded on June 05,2021 at 12:57:09 UTC from IEEE Xplore. Restrictions apply. low-cost UWB (ultra-wideband) modules. The shared velocity information and individual range measurements are used to estimate the relative poses among robots. The uncertainties in pose estimation are taken into account for the VO based collision avoidance algorithm. The main contributions of this work include: 1) Developing a distributed collision avoidance architecture, in which each robot performs its own motion planning and shares the velocity information with each other. Such a distributed framework is robust against local failures and enables large-scale robot groups. 2) Developing a particle filter based algorithm to estimate the relative poses among robots from the range only measurements and shared information on velocities. 3) Developing an extended reciprocal collision avoidance algorithm to tackle the uncertainties in the estimated poses guarantees the robust performance of collision avoidances in the real-world applications. 4) Developing both simulation and experiment platforms for large-scale robot group navigation. The proposed approach is validated using multiple Turtlebots equipped with UWB modules in comparison with other state-ofthe-art methods. The rest of the paper is organized as follows. Section II discusses the related approaches to multi-robot collision avoidance. Section III describes the system setup and problem statement. Section IV presents the proposed approach. Section V provides simulation and experiment results. Section VI concludes the paper and outlines future work."
        ]
    },
    {
        "image_filename": "designv11_63_0000703_s12206-020-0122-7-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000703_s12206-020-0122-7-Figure12-1.png",
        "caption": "Fig. 12. Bolt and nut with multiple specimens.",
        "texts": [
            " The first step is when the load acts on the fastening component that is in contact with the bolt. The second stage is the case where the load acts on the fastening component without contact with the bolt. If loosening occurs in the first condition, it is a direct cause of the slip that happens at the contact between the bolt and the fastening component. These results are not different from those of simple structures. However, in the case of loosening occurring under the second condition, looseness occurs through the two-stage process described above. As shown in Fig. 12, it is possible to identify the applied lateral load. If the magnitude of the applied lateral load is greater than the frictional force (\u2460 - \u2461, \u2461 - \u2462 generated from the two parts) between the fastening components caused by the tightening force of the bolt. Then the fastening component is slipped. In this case, slip is not generated in other fastening components, and alone movement is generated only in fastening component that has been subjected to a load. This is defined as the primary loosening",
            " All parts are configured with solid elements, and the 8-node hexahedron elements with reduced integration (C3D8R) were adopted in all specimen models and 4-node tetrahedron elements (C3D4) were applied in all the bolts and nuts model because of their complicated shapes. Fig. 16 shows the boundary condition for tightening analysis of the bolt. The both ends of the specimen are constrained to simulate the fixing of the specimen during the tightening analysis. The contact conditions are a total of 5 parts as shown in Fig. 12; Bolt-fastening component 1, fastening component 1 - fastening component 2, fastening component 2 - fastening component 3, fastening component 3 - nut, bolt - nut. The coefficient of friction was obtained by the data derived from the test. Tightening torque of the same magnitude as the test is applied to the bolt head. In addition, the nut is constrained by all the degrees of freedom except for the axial length of the bolt. At this time, the fastening force is derived through the contact force among the fastening components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002338_012093-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002338_012093-Figure1-1.png",
        "caption": "Figure 1. In this figure is show a helicopter with all of it elements. Our case of study is the roof structure focused on internal forces [1]",
        "texts": [],
        "surrounding_texts": [
            "Content from this work may be used under the terms of the Creative Commons Attribution 3.0 licence. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI.\nPublished under licence by IOP Publishing Ltd\n1. General knowledge of structure elements in local and global coordinates",
            "10th EASN 2020 IOP Conf. Series: Materials Science and Engineering 1024 (2021) 012093\nIOP Publishing doi:10.1088/1757-899X/1024/1/012093\nStatic indeterminate structures consist of branches and nodes. Branches can be divided into even smaller non-uniform parts according to the external load, into elements.\nThe node the place where the elements come together and can form nodes with two, three or more elements. If we know the displacements of the nodes or the factors inside the nodes as well as the load on them, all the factors in the complete can be determined. A particular element can be studied in many ways. The dependencies between the joint force factors with the joint displacements can be determined. Each element of a structure has two nodes. Each node actually represents a cross section. The cross section in the plan performs three movements, horizontally, vertically, and rotationally. These movements also represent joint degree of freedom. If we had horizontal, vertical, and internal force, these factors would do work, so any movement (displacement) is linked to a force factor when we use energy theorems.\nIn general structure all the elements are connected with each other according to the principle of construction, where the continuity of the deformation and the balance of the joints in the elements are fulfilled.\nExternal loads may be in the structure at different locations and may not match the nodes of the elements. Despite the various maneuvers we can do, at the distributed load must be equivalent to the joints, because at the joints we have displacements. Equivalence will be done according to potential energy or internal work, so the work of the forces in the scheme is equal to that of the equivalent forces in the joints. A simple method of determining equivalent forces is to find the reactions of the particular element scheme."
        ]
    },
    {
        "image_filename": "designv11_63_0002979_s10514-021-09996-3-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002979_s10514-021-09996-3-Figure13-1.png",
        "caption": "Fig. 13 Left: The planned contact transition sequence using PFS Only. Right: The planned contact transition sequence using the proposed framework. Left and right palm contact are shown in red and green, respectively",
        "texts": [
            " For each torso translation, we collect data over sampled randomlytilted surface environments, and train a traversability estimator with the mobile manipulator\u2019s contact transition model. We also collected a motion plan library for the mobile manipulator. However, there are abundant contacts in the test environment. Therefore, in this test, the proposed framework outputs a torso pose guiding path with only one segment, and uses PFS (which accounts for traversability) for that segment. The result in Fig. 13 shows that the proposed framework has a much shorter planning time, but the resulting path takes more steps. Since the PFS Only planner does not consider traversability, the planner will explore the slightly-shorter path above the window first. However, the gap created by the pipe makes the path above the window require more steps than the path under the window and the PFS Only planner, misled by the heuristic, spends a large amount of time rejecting states around the gap before searching below thewindow"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001505_icecce49384.2020.9179221-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001505_icecce49384.2020.9179221-Figure3-1.png",
        "caption": "Fig. 3. ADAMS model for the 4-bar mechanism",
        "texts": [
            " ADAMS/View is a graphical user environment that allows users to build motion models, simulate them with ADAMS, and examine a variety of results. Virtual prototyping in ADAMS is executed in three steps, i.e., i) build the model (using links, joints, contacts, torques, etc.); ii) test the model (using simulations and plots); iii) model validation (by comparing the results with published literature). The 4-bar mechanism is created in ADAMS/View by following the various steps discussed above. The final 4-bar mechanism is shown in Fig. 3. ADAMS simulations were performed for the 4-bar mechanism with the parameters shown in Table 1. VI. RESULTS AND DISCUSSIONS After creating the 4-bar mechanism in ADAMS/View, the next step is to simulate this model and to get the output results. The results are obtained by running the simulation over time intervals of 0-10 seconds in ADAMS. The output results for the position, velocity and acceleration analysis for four-bar mechanism are discussed below. A. Position Analysis The torque is applied to the input link"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001884_1350650120975519-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001884_1350650120975519-Figure2-1.png",
        "caption": "Figure 2. Contact between the inner and the outer rings under normal load.",
        "texts": [
            " It is noteworthy that the interval DR of contact surfaces has a significant effect on the distribution of contact pressure.18 When the interval is considerably small, the contact pressure concentrates near the edges, moving towards the center with the increase of DR. In this study, the interval is set to be zero to study the stress concentration on the free boundary, which means that the radius of both contact surfaces are equal. According to the conformal contact theory,15 the circumferential contact between the inner and the outer rings under normal load is an axis-symmetric elastic issue as shown in Figure 2. Therefore, the contact pressure is supposed to obey an axis-symmetric distribution in the following way p r\u00f0 \u00de \u00bc p0 1 r2 a2 n (1) where p0 is the maximum contact pressure, r is the distance away from the contact center, a is the radius of contact area and n is a constant index related to the contact condition. It is noted that the index n physically describes and determines the uniformity of contact pressure distribution. The greater n is, the more concentrated the distribution, while the smaller n is, the more dispersed the distribution",
            "3 for studying the effects of friction on the contact pressure, while those of the other two contacts are fixed as 0.3 to avoid any possible slipping or rotation. All displacements of the base\u2019s side faces are prohibited to restrain the movement of the whole system with no constraint imposed on the end faces of the outer ring. All the materials in this model are chosen as steel with Young\u2019s modulus E\u00bc 210GPa and Poisson ratio \u00bc 0:33. A uniformly distributed pressure P is implemented on the dabber as shown in Figure 2 to exert radial load to the spherical plain bearing. Two paths are selected along axial and circumferential directions on the internal surface of the outer ring to display the contact pressure distribution, as shown in Figure 5, where Path 1 runs through the axial width of the outer ring, and Path 2 covers the half-circle of the internal surface. y z\u00f0 \u00de u z\u00f0 \u00de M z\u00f0 \u00de Q\u00f0z\u00de 8>>< >>: 9>>= >>; \u00bc A1 az\u00f0 \u00de A2 az\u00f0 \u00de A3 az\u00f0 \u00de A4 az\u00f0 \u00de 4a A4 az\u00f0 \u00de a A1 az\u00f0 \u00de a A2 az\u00f0 \u00de a A3 az\u00f0 \u00de EI 4a2A3 az\u00f0 \u00de EI 4a2A4 az\u00f0 \u00de EI a2A1 az\u00f0 \u00de EI a2A2 az\u00f0 \u00de EI 4a3A2 az\u00f0 \u00de EI 4a3A3 az\u00f0 \u00de EI 4a3A4 az\u00f0 \u00de EI a3A1 az\u00f0 \u00de 2 66664 3 77775 y0 u0 a M a2 EI Q a3 EI 8>>>>< >>>>: 9>>>>= >>>>; (2) The whole model is meshed into tetrahedron with a type of C3D10M and an initial global seed size of 5, followed by an adaptive process of four iterations to validate mesh convergence under the maximum pressure of 100MPa and l \u00bc 0:3",
            " Considering that the displacement in y direction is negative, the curves in Figure 11 indicate a decreasing tendency of the downward distance with the increase of l, which is in turn accompanied by the increasing contact area in the central region as discussed above. Table 3 describes the variation of maximum contact radius under different loads and friction coefficients. Since the circumferential contact between the inner and the outer rings under normal load basically follows the conformal contact theory, the contact radius is actually determined by the geometrical shape of the contact area. In this study, the curvatures of both contact surfaces are set strictly equal, inferring a spherical contact area between the inner and the outer rings as shown in Figure 2. Therefore, it is predictable that the contact radius approximately equals to half of the outer circumferential spherical diameter, i.e. a dk=2 . Moreover, it can be seen from Figures 10 and 11 that the deformation of the contact surfaces is actually several orders of magnitude smaller than the contact radius even under very large normal loads, which indicates that the alteration of the contact radius with the load and friction coefficient can be very small as validated by Table 3. The relationship of n and the load and friction coefficient is presented in Table 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000503_012020-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000503_012020-Figure4-1.png",
        "caption": "Figure 4. Initial water flow in the waterwheel",
        "texts": [],
        "surrounding_texts": [
            "The phenomenon of water flow in the waterwheel has been studied using CFD simulation using Ansys Student Version 19.1 calculated based on flow assumptions, with an initial flow velocity of 5.00 m/s by entering the gravity acceleration value 9.81 m/s2. The position of the water wheel is assumed to be half float. The model of approach used is k-epsilon. By using water-liquid fluid material, stating that the waterwheel moves based on the flow of water by 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 selecting \"Relative to cell zone: fluid-solid\". The speed of water produced around the mill is 2.19 m/s while the speed of water after passing the water mill is 5.07m/s with the pressure experienced by the waterwheel is 42326 Pa."
        ]
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure19-1.png",
        "caption": "Fig. 19. UMF diagrams in STPSPM machines under rotating eccentricity. (a) UMF distribution, 10-pole/12-slot. (b) UMF components, 10-pole/12-slot. (c) UMF distribution, 10-pole/9-slot. (d) UMF components, 10-pole/9-slot.",
        "texts": [
            " (a) (b) (c) (d) For the 10-pole/9-slot STPSPM machine, the variation of post-demagnetization UMF with static eccentricity shows similar rules with the 10-pole/12-slot STPSPM machine, as shown in Fig. 17. However, when the rotating eccentricity is applied, the 1st order harmonic in Fx and Fy decreases after demagnetization, which is opposite to the rules in the 10- pole/12-slot STPSPM machines. The reasons for such difference are analyzed as follows. The UMF distribution for the symmetrical SPM machine on the rotor at a specific rotor position is shown in Fig. 19(a). The total UMF for the symmetrical SPM machine can be simply regarded as the vector sum of F1 and F2, which are the magnetic forces generated by the half machines, as shown in Fig. 19(b). When the eccentricity is not considered, F1 equals to F2 and thus the total UMF is 0N. When the rotating eccentricity is applied, F1 increases to F1-eccen due to smaller airgap whilst F2 decreases to F2-eccen due to larger airgap. Therefore, the total UMF is no longer 0N under rotating eccentricity. Once demagnetization happens due to 3PSC, F1eccen and F2-eccen are reduced to F1-eccen-pd and F2-eccen-pd, respectively. However, since the demagnetization is severer for the rotor part facing larger airgap, F2-eccen drops more than F1eccen and thus the total UMF increases after demagnetization. On the contrary, for the asymmetric SPM machines, the UMF is considered as a whole and thus decreases after demagnetization, as shown in Fig. 19(c) and (d). -400 -200 0 200 400 0 60 120 180 240 300 360 F x (N ) Rotor position (Mech. Deg.) Fx-before Fx-after 0 100 200 300 400 0 1 2 3 4 5 6 7 8 9 10 11 12 13 F x (N ) Harmonic order 0 16 11 12 13 Fx-before Fx-after zoom -400 -200 0 200 400 0 60 120 180 240 300 360 F y (N ) Rotor position (Mech. Deg.) Fy-before Fy-after 0 100 200 300 400 0 1 2 3 4 5 6 7 8 9 10 11 12 13 F y (N ) Harmonic order Fy-before Fy-after 0 16 11 12 13 zoom Authorized licensed use limited to: California State University Fresno"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure16-1.png",
        "caption": "Figure 16. Drawing of the exterior face of the car door.",
        "texts": [],
        "surrounding_texts": [
            "The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article."
        ]
    },
    {
        "image_filename": "designv11_63_0002822_0309524x211015271-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002822_0309524x211015271-Figure2-1.png",
        "caption": "Figure 2. Schematic of the test article (2.3-MW drivetrain).",
        "texts": [
            " This test bench is part of the Energy Innovation Center at Clemson University that also includes a 15-MW test bench and a 15-MW grid emulator (Schkoda and Fox, 2014). Figure 1 depicts the 2.3-MW drivetrain installed on the 7.5-MW test bench. This figure shows the main components of this test bench, which are an 8.5-MW asynchronous water-cooled motor (blue), an adaptation gearbox having a ratio of 100:1, a low-speed coupling, a load application unit (LAU), and an adapter to connect the LAU to the drivetrain under test (test article). A schematic of the test article is shown in Figure 2. The 2.3-MW drivetrain design has a single main bearing, multi-stage planetary/helical gearbox, doubly-fed induction generator, and a high-speed shaft coupling. The maximum torque and speed of the test bench are 6.5 MNm and 20 rpm, respectively. Other features of the test bench are an isolated foundation and adjustments to accommodate drivetrain tilt angles of 4 \u20136 . The LAU of this test bench consists of 24 individual hydraulic actuators (red cylinders in Figure 1) capable of generating radial and axial (thrust) forces of 62 MN and bending moments of 610 MNm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001229_15397734.2020.1787841-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001229_15397734.2020.1787841-Figure9-1.png",
        "caption": "Figure 9. Secondary axes graph for locating the rotation axis 2u6:",
        "texts": [
            " (100), and the corresponding vectors are made unitary to obtain two possible values of 1u3: 1u3a \u00bc 19 ffiffi 2 p 95143\u00fe 478344 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1268455324557594 134825216975072 ffiffi 3 pp 2 ffiffi 2 p 1020077\u00fe 2457452 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1268455324557594 134825216975072 ffiffi 3 pp 4 ffiffi 2 p 1020077\u00fe 2457452 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1268455324557594 134825216975072 ffiffi 3 pp 2 6666666664 3 7777777775 , 1u3b \u00bc 33 ffiffi 2 p 18422\u00fe 4475 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3200988920512 1542464253205 ffiffi 3 pp 1 2 ffiffi 2 p 494129\u00fe 478969 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3200988920512 1542464253205 ffiffi 3 pp ffiffi 2 p 494129\u00fe 478969 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3200988920512 1542464253205 ffiffi 3 pp 2 6666666664 3 7777777775 : (127) Furthermore, with the two values of 1u3, using the process outlined in this section the two possible values of 2u6a and 2u6b are attained Summarizing, in the first part of this problem, the secondary rotation axis 2u6 was chosen as known, and the secondary rotation axis 1u5 was the overconstrained rotation axis. The final results given by Eq. (99), provide the secondary rotation axis 2u6: In the second part of this problem, the secondary rotation axis 1u3 was chosen as known, obtaining the results given by Eq. (128), that also provided the secondary rotation axis 2u6, see Fig. 9. Comparing the results, the only coincident result is given by 2u6b \u00bc 2 3 3075897\u00fe 2184428 ffiffi 3 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 42605013632135\u00fe 22945428529226 ffiffi 3 pp 1 3 ffiffi 3 p 759823\u00fe 2038007 ffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 42605013632135\u00fe 22945428529226 ffiffi 3 pp 1 3 6192259\u00fe 3861410 ffiffi 3 p ffiffi 3 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 42605013632135\u00fe 22945428529226 ffiffi 3 pp 2 66666666664 3 77777777775 \u00bc 4829412908\u00fe 6108507066 ffiffi 3 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1592101255916815155465 844010536641692794038 ffiffi 3 pp 14181374901 ffiffi 3 p \u00fe 27677483191ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1592101255916815155465 844010536641692794038 ffiffi 3 pp 9351961993ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1592101255916815155465 844010536641692794038 ffiffi 3 pp 2 6666666664 3 7777777775 \u00bc2u6b It can be proved that both fractional and irrational expressions are equivalent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000270_s40815-019-00760-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000270_s40815-019-00760-5-Figure1-1.png",
        "caption": "Fig. 1 Truck\u2013Trailer Model",
        "texts": [
            " Remark 2 Generally, it is noticeable that the finite-time stability can be derived from the finite-time boundedness by assuming the values of external noises and the forces of controller as null. Specifically, in Lyapunov based asymptotically stable, the occurrence of stability may or may not be within a finite-time duration. So, the finite-time stability is more suitable one compared with Lyapunov asymptotic stable for the practical models. In this section, an illustration of computer simulated trucktrailer model shown in Fig. 1 is considered to show the effectiveness of our proposed controller. Consider the following computer simulated truck\u2013trailer model: dx1\u00f0t\u00de \u00bc v t Lt0 x1\u00f0t\u00de \u00fe v t lt0 u\u00f0t\u00de; dx2\u00f0t\u00de \u00bc v t Lt0 x1; dx3\u00f0t\u00de \u00bc v t t0 sin x2 \u00fe v t 2L x1 ; where the model parameters are known as l \u00bc 2:8, L \u00bc 5:5, v \u00bc 1, t \u00bc 2, t0 \u00bc 0:5. Let h\u00f0t\u00de \u00bc x2 \u00fe v t 2L x1 and d \u00bc 10 t0=p be taken as premises variable and the following fuzzy rules are considered to represent the truck\u2013trailer model with external disturbance and time-varying actuatorfault: Rule 1 If h\u00f0t\u00de is about 0 rad, then dx\u00f0t\u00de \u00bc A1x\u00f0t\u00de \u00fe B1 U\u00f0t\u00de \u00fe F\u00f0t\u00de\u00f0 \u00de \u00fe D1d\u00f0t\u00de Rule 2 If h\u00f0t\u00de is about p rad to p rad, then dx\u00f0t\u00de \u00bc A2x\u00f0t\u00de \u00fe B2 U\u00f0t\u00de \u00fe F\u00f0t\u00de\u00f0 \u00de \u00fe D2d\u00f0t\u00de, where the system parameters are known as A1 \u00bc v t Lt0 0 0 v t Lt0 0 0 v2 t2 2Lt0 v t t0 0 2 6 6 6 6 6 6 4 3 7 7 7 7 7 7 5 , A2 \u00bc v t Lt0 0 0 v t Lt0 0 0 dv2 t2 2Lt0 dv t t0 0 2 6 6 6 6 6 6 4 3 7 7 7 7 7 7 5 , B2 \u00bc B1 \u00bc v t lt0 0 0 2 6 4 3 7 5, D1 \u00bc D2 \u00bc 0 0 0:3v t 4t0 2 6 4 3 7 5, r1\u00f0h\u00f0t\u00de\u00de \u00bc 1 1 1\u00feexp 3\u00f0h\u00f0t\u00de 0:5p\u00de\u00f0 \u00de 1 1\u00feexp 3\u00f0h\u00f0t\u00de\u00fe0:5p\u00de\u00f0 \u00de , r2\u00f0h\u00f0t\u00de\u00de \u00bc 1 r1\u00f0h\u00f0t\u00de\u00de"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001894_icem49940.2020.9270713-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001894_icem49940.2020.9270713-Figure13-1.png",
        "caption": "Fig. 13. Stress distribution by the analysis of the shrink fitting",
        "texts": [
            " The minor loops caused by the slot harmonics have DC components. Using the hysteresis Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 24,2020 at 06:42:49 UTC from IEEE Xplore. Restrictions apply. model such as Play model, it is possible to estimate the minor loops, including a DC offset, accurately. The loss due to the stress is observed mainly in the low load region. To consider the effect of the stress, a stress analysis is conducted to obtain the stress distribution in the machine as shown in Fig. 13. A high-stress distribution is observed in the root of the stator tooth where the change in the material\u2019s magnetic property will be the largest. In the magnetic analysis, a stress-dependent iron loss property is used as shown in Fig. 14. The increased ratio of the iron loss is larger for low magnetic fields and this is the reason why the effect of the stress is the highest at the low load region. The resulting iron loss density distribution is shown in Fig. 15. The accuracy of the FEA-based efficiency map was evaluated, comparing it to measurement results for an IPM Authorized licensed use limited to: Auckland University of Technology"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure6-1.png",
        "caption": "Figure 6. The machining model of the hourglass worm in the VERICUT. (a) front view, (b) side view, and (c) a sectional view.",
        "texts": [
            " When the B-axis rotation center is at the ps point, the initial state of the process system is represented by [xs, zs, cs, bs]. When the center of rotation of the B-axis is at point pe, the state of the process system is represented by [xs, zs, cs, bs]. According to Table 1, Figure 5, and the tool setting operation, the above coordinate values are easy to determine. Based on the Fanuc oi CNC system, the core G code for simulation are shown below: G00 Xxs Zzs Ccs Bbs; G18; G02 Xxe Zze Cce Bbe Rrc1; M30; The machining model of the hourglass worm is shown in Figure 6. Since the tool refers to a finger tapered tool, the virtual machining model is a finger tapered surface enveloping hourglass worm according to the meshing principle. As shown in Figure 6, the virtual model is a right-handed multi-threads hourglass worm. Through the simulation processing, the corresponding hourglass worm is obtained. The CZXB function of the machine tool is verified by simulation processing. The simulation results show that the CZXB method is completely consistent with the theoretical analysis and can be applied to actual production. To verify the flexibility of the CZXB method, a trial manufacturing of an hourglass worm is carried out. A trial manufacturing of the hourglass worm is carried out for further verification of the flexibility of CZXB method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001054_tcns.2020.2996938-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001054_tcns.2020.2996938-Figure1-1.png",
        "caption": "Fig. 1. The configuration of a robotic arm.",
        "texts": [
            " Suppose the dynamic controller (A\u03c7,i, B\u03c7,i, C\u03c7,i) solves the H\u221e control problem for the system A\u0304i R\u0304i B\u0304i C\u0304i 0 0 C\u0304i 0 0  and the H\u221e norm of the closed-loop system is less than\u221a \u03b1$/(N\u03b2$) + co for some co > 0. There exist neighborhoods Wi of wi = 0, i = 1, \u00b7 \u00b7 \u00b7 , N , on which the robust output synchronization problem is solved by a distributed controller of the form ui = C\u03c7,i\u03c7i + \u03a8iT \u22121 i \u03b7i \u03c7\u0307i = A\u03c7,i\u03c7i +B\u03c7,i(yi \u2212 Covi) v\u0307i = Aovi +BoC\u03b6\u03b6i \u03b6\u0307i = A\u03b6\u03b6i +B\u03b6 \u2211 j\u2208Ni aij(yj \u2212 yi), \u03b7\u0307i = Mi\u03b7i +Niui + Ti\u03a5iBoC\u03b6\u03b6i \u2212NiB+ i XiBoC\u03b6\u03b6i. (39) Proof: Bearing Proposition 2.1 in mind, the proof is similar to that of Theorem 4.3. Consider a group of manipulators each of which has the configuration shown in Fig. 1. Suppose the ith robotic arm has mass Mi and length Li, and the gripper can be regarded as a mass point with mass mi. Denote by ri the position of the gripper and \u03b8i the angle of the arm, equations for such a system are (see [32]) (mir 2 i +MiL 2 i /3)\u03b8\u0308i + 2mirir\u0307i\u03b8\u0307i = \u03c4i mir\u0308i \u2212miri\u03b8\u0307i 2 = Fi, i = 1, \u00b7 \u00b7 \u00b7 , N (40) where Fi and \u03c4i are the external torques. The objective is to synchronize the position of gripper and the angle of the arm denoted by yi := [\u03b8i, ri] towards constants, i.e., limt\u2192\u221e yi(t) = [\u03b8o, ro]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.5-1.png",
        "caption": "Fig. 57.5 Boundary conditions in FEA model",
        "texts": [
            "4a) was used for the rider seating region of the component, and seed size of 3 mm was used for rubber cushion mounting holes and the guide for rubber cushion (Fig. 57.4b). This seed size was varied depending on the thickness of the component. In order to arrive at the boundary conditions, a seat load impression test was carried out using blue print method on existing seat base under the load conditions of: nopassenger, solo passenger, and twin passenger configurations. The boundary conditions were applied at the initial step in the CAD model (Fig. 57.5). The seat base bottom rubber cushion mounting and the rear rubber grommet mounting region were constrained along the vertical direction. A uniform pressure distribution of 18.5 kPa (corresponding to 2600 N of distributed load) was applied over the rider and the pillion seating area in the vertical direction (Fig. 57.6). The failure criteria used for homogeneous materials are not sufficient for predicting failure in composite lamina. This is because the planes along which the lamina may be possibly be the weakest need not be the direction of principal stresses in a lamina"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure6-1.png",
        "caption": "Fig. 6. The shape plot of temperature distribution at the end of the Laser beam.",
        "texts": [],
        "surrounding_texts": [
            "From the experimental results and the discussions, the following conclusions can be drawn: i. The deposited clad of the composites varied with the change in the selected process parameters, this in turn affected the physical, mechanical, and metallurgical properties of the composites. ii. From the analysis of the results obtained from characterization, the combination of laser power of 900 and 1000 W, powder feed rate of 2.5 g/min, gas feed rate of 2.0 L/min, spot diameter of 2 mm and scanning speed of 0.8 and 1.0 m/ min produced composites with an increased percentage of dilution rate, aspect ratio and enhanced powder efficiency. Based on the analysis of these findings, these process parameters are optimum for producing highly reinforced composites. iii. The geometrical characteristics of the composites were influenced by the laser power, scan velocity and the powder feed rate. Finally, from the computational modelling, the movement of laser from the starting point to the end changes the temperature distribution and enhanced microstructures in the melt pool are depended on the laser input, scan velocity and the faster cooling rate. CRediT authorship contribution statement Dr. Olawale S. Fatoba is the co-supervisor of the Master student that work on this work and he played a crucial role in the modelling and thorough supervision of the work. Mr. Adedoyin M. Lasisi is the Master\u2019s student that worked on this project. He carried out characterizations and experimental procedures of the work. He writes the literature and did carry out the validation of the work. Dr. Omolayo M. Ikumapayi is a project administration, he helps in the drafting of the literature and in put the entire work in a right journal formal and revising the work. He revised and did the editing of the work. Dr. Stephen A. Akinlabi helps in proof-reading the entire manuscript and also helps in reviewing the work and contributes immensely in the methodology and discussions of the results. He also provides additive materials. Prof. Esther T. Akinlabi happens to be the main supervisor of the Master\u2019s student that did this work. She provides funding for materials and software and She conceptualized the project and did supervision for the success of the research. 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."
        ]
    },
    {
        "image_filename": "designv11_63_0001946_s10778-020-01029-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001946_s10778-020-01029-3-Figure1-1.png",
        "caption": "Fig. 1",
        "texts": [
            " The problem of determining the coefficients of the optimal controller for stabilizing the system was solved in [9] using the particle swarm methods and genetic algorithm. From the point of view of the time required to solve the problem, the first method turned out to be more effective. In this study, we solve the problem of optimal control of the stabilization of an inverted pendulum with a flywheel. In the problem statement, an integral criterion and a control constraint are used, and the solution to the problem is obtained in the form of program control and feedback. 1. Flywheel Inverted PendulumModel.Consider a systemwhose dynamic model is shown in Fig. 1. In this study, we assume a small deviation of the pendulum from the vertical. This assumption is based on the fact that the dynamic system shown in Fig. 1 can be transferred to the vicinity of the equilibrium position (vertical position of the pendulum). For example, if the initial position of the system is described by the conditions ( ) , ( ) , ( )0 0 0 0 0 , then it can be transferred from this position by energy pumping [13]. 462 1063-7095/20/5604-0462 \u00a92020 Springer Science+Business Media, LLC 1 National University of Life and Environmental Sciences of Ukraine, 12v Geroev Oborony St., Kyiv, Ukraine; e-mail: romasevichyuriy@ukr.net. 2 S. P. Timoshenko Institute ofMechanics,NationalAcademyof Sciences ofUkraine, 3NesterovaSt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000380_0954406219897688-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000380_0954406219897688-Figure12-1.png",
        "caption": "Figure 12. Photograph of the experimental platform.",
        "texts": [
            " The gear ratio of the two-stage gearbox is 3.59, with 23/39 for the first stage, and 25/53 for the second stage. A small spalling fault was seeded in the second-stage gear pair as shown in Figure 9, the depth (ds), width (ws) and length (ls) of the spalling fault are 1mm, 3mm and 15mm, respectively. Four accelerometers (Kistler 8203A50) are mounted on the bearing casing of the gearbox, an LMS (SC305-VTP/3P92-B) data acquisition system (DAQ) and a PC were used to collect the vibration data for further processing. Figure 12 is the photograph of the data collection process. The sampling frequency is 5120Hz. It is noted vibration signal of the accelerometer mounted on the bearing casing of the second stage is analyzed in this study. In order to validate the effectiveness of the proposed VM2RL method, the vibration signal acquired in 700 r/min and 60Nm condition with 5120 points was analyzed as an example. The rotation frequency fr is calculated as 6.88Hz. The transient impulses were not very clear due to the background noise masking in both the waveform in Figure 13(a) and envelope in Figure 13(c)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003208_s42835-021-00842-1-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003208_s42835-021-00842-1-Figure3-1.png",
        "caption": "Fig. 3 Duration of data collection for learning and control: N1 gait cycles for collecting data using a simple master\u2013slave position control; N2 gait cycles driven by MLPC",
        "texts": [
            " Taking only human walking into account, although the human gait is different within one person over time, this difference is not so significant and still ensures that the desired trajectory for prediction will be inside the regions that it has been learned. It means that movement data collected by that way is able to capture the nonlinear dynamics of the exoskeleton and the corresponding human\u2013robot (16)Dk = MT k Mk, with Mn+1 k = Mn k \u2212 Jk Mk 1 3 interaction torques. These data, naturally, can be collected by controlling the exoskeleton to track the human motion with a non-model based controller, say a master\u2013slave position controller. As shown in Fig.\u00a03, it is assumed that number (N1 + N2) iterations of the walking gait cycle (only considered in swing phase) will be performed. The exoskeleton is controlled by a simple master\u2013slave PD controller with high gains in the first N1 iterations of the trajectory and then switched to the partitioned controller with the model being learnt. Half of the observed data was used for training the model and half for testing and estimating the accuracy of the learned model. In order to obtain this adequate movement data, the design of the position master\u2013slave controller for collecting data depends on several setup issues, for example, a proper mechanical connection between the master (human) and the slave (exoskeleton) as exampled in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000691_s11837-020-04089-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000691_s11837-020-04089-5-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the EBCHM.",
        "texts": [
            " Amush Mushy zone constant b Liquid fraction cp Specific heat (J/kg K) g Acceleration due to gravity (m/s2) H Enthalpy (J/kg) h Sensible enthalpy (J/kg) href Reference enthalpy (J/kg) DHf Pure solvent melting heat (J/kg) k Thermal conductivity (K/m K) L Characteristic length (m) P Pressure (Pa) q Density (kg/m3) s Stress tensor (N/m2) Tref Reference temperature (K) T Temperature (K) t Time (s) ~v Cell velocity (m/s) ~vp Pull velocity (m/s) In recent years, electron beam cold hearth melting (EBCHM), the schematic diagram of which is shown in Fig. 1, is playing a major role in the manufacturing of high-performance titanium products.1\u20133 Because of the design of the cold hearth, EBCHM has an outstanding performance in the inclusions control compared with vacuum arc remelting. Additionally, a strip coil can be rolled out from slab ingots produced by EBCHM without the need for forging and stamping, which significantly reduces the cost of many titanium products.4 One of the characteristics of EBCHM is high energy consumption. During the 10\u201315 h casting process for 8-m to 10-m ingots, high-energy intensity electron beams are used as the heat source for melting feedstocks and maintaining the fluidity of the molten alloy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002113_0954407020984668-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002113_0954407020984668-Figure2-1.png",
        "caption": "Figure 2. ESF friction model (from Pfeffer14,15).",
        "texts": [
            " Referring to a modeling approach proposed by Pfeffer,14,15 the analytical model more appropriate and representative of the friction, produced by connecting constraints between the steering column and the vehicle body, is that called Exponential Friction Spring (ESF) model (1): FESF= Flim 1 e fESF x sign _x\u00f0 \u00de sign _x\u00f0 \u00de x= x xR + xB \u00f01\u00de with FESF exponential spring friction Force, xR displacement at the instant of motion reversal, xB off set to be added to the displacement so that the branch of increasing and decreasing exponential behavior maintains continuity, fESF = kESF/Flim a variable representing the stiffness of the exponential spring, Flim a saturation force value and kESF the slope of the function at force zero value. Figure 2 schematically represents this model (from Pfeffer14,15) Instead, to correctly represent the sliding of the rack to the support, the model considered as a reference is that defined as Exponential Spring Friction with Maxwell Element (ESF-M) (2): FESF M =FESF+FM FM = kM x2 x1\u00f0 \u00de=CM _x1 ) FM = kM x2 \u00f0t2 t1 FM CM dt 0 @ 1 A \u00f02\u00de Figure 3 represents ESF-M model (from Pfeffer et al.14). In equation (2) FM is Maxwell element force (see Figure 4), x1, _x1 the state variables of displacement and velocity related to the degree of freedom connected to the damper, x2 the displacement state variable corresponding to the degree of freedom connected to the spring, kM spring stiffness, CM damper damping, t1, t2 the time instants that define the friction force evaluation interval"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000089_ev.2019.8892977-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000089_ev.2019.8892977-Figure4-1.png",
        "caption": "Fig. 4. PMVM-BM magnetic flux density map and flux lines distribution",
        "texts": [],
        "surrounding_texts": [
            "In this section it is presented the magnetization of the cores and the flux lines passing through inner stator, the inner air gap, the PM in the rotor, the rotor core, the outer air gap, the PMs in the outer stator and the outer stator core. From Figure 3 to Figure 5, a) can be highlighted that the magnetic flux density does not exceed 2.3 T, although PMVM-SMMRS rotor and outer stator is more saturated in comparison with PMVM-BM and PMVM-SMM. The flux lines distribution is represented in Figure 3 to Figure 5, b) where the role of PM can be highlighted. Not all the flux lines produced in the stator secondary poles are closing through the outer stator which results in the influence of the coercive magnetic field produced by the PM in the outer stator."
        ]
    },
    {
        "image_filename": "designv11_63_0002822_0309524x211015271-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002822_0309524x211015271-Figure12-1.png",
        "caption": "Figure 12. Top and side views of the 2.3-MW drivetrain MBS model.",
        "texts": [
            " Measured bending moments and speed and commanded forces The comparison of the simulation results using the input sets A and B targets the question of the level of tracking error being acceptable or not whereas input data sets B and C target the cross-coupling effect. The MBS model of the 2.3-MW drivetrain tested on Clemson University 7.5-MW test bench was developed at the Energy Innovation Center of Clemson University using inputs from GE Renewable Energy. Top and side views of the model are presented in Figure 12. The main shaft is shown in cyan. The point of application of the loads and speed inputs to the drivetrain is indicated with a red circle. The two circles with doted purple lines capture the main bearing to the left of both views and the high-speed shaft and its coupling to the right. The bedplate, generator frame and the support tower connecting the drivetrain to the ground of the test bench are shown in grey. The generator frame is to the right of both views. The first stage of the gearbox is shown in blue and the rest of the gearbox is shown in orange"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000555_j.jmapro.2020.01.054-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000555_j.jmapro.2020.01.054-Figure2-1.png",
        "caption": "Fig. 2. Schematic of an aircraft landing gear fabricated by forging and additive manufacturing.",
        "texts": [
            " Integrating the forging technology with the AM technology would combine their advantages. First, the forging can ensure that the size, structure and performance of the main part of the compositely formed component are stable and controllable, and meet the needs of mass production. Second, the AM technology can ensure the manufacturing convenience of complex structures in the component, improve material utilization and save cost. For example, for an aircraft landing gear with complex budlike structures (see Fig. 2) which used to be shaped by forging, if its simple substrate (the gray https://doi.org/10.1016/j.jmapro.2020.01.054 Received 11 June 2019; Received in revised form 5 December 2019; Accepted 29 January 2020 \u204e Corresponding author. E-mail address: huangliang@hust.edu.cn (L. Huang). Journal of Manufacturing Processes 52 (2020) 79\u201395 1526-6125/ \u00a9 2020 The Society of Manufacturing Engineers. Published by Elsevier Ltd. All rights reserved. T portion shown in Fig. 2) is forged and its partially complex structures (the yellow portions shown in Fig. 2) on the substrate are fabricated by the AM technique, it will be easier and cheaper to fabricate such complex budlike structures while the forged substrate has a controllable size, structure and performance. In addition, material utilization is improved owing to the application of the AM technique. For compositely formed components fabricated by forging and AM, the substrate geometry is an important factor for determining the structure and location of the bonding zone between the substrate and AM part, which considerably affects the temperature and stress distribution during the subsequent AM process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001199_032121-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001199_032121-Figure1-1.png",
        "caption": "Figure 1. The proposed sandwich structure.",
        "texts": [
            " In recent years, many lattice structures with 3D lattice structures have been designed. This project aims to find has the light quality high strength characteristics of dot matrix composites reinforced polymer structure, using modelling software to design a lattice structure and a series of mechanical properties of the structure analysis, considering the lattice composites reinforced polymer mechanics performance optimization, study the effect of lattice structure of composite sandwich structure are studied. In this study, we designed a lattice sandwich structure as shown in figure 1. The sandwich structure consists of two aluminum face sheets and a lattice core. The core is a composite structure composed of body core structure and face core structure. Its overall size is 160 mm \u00d7 40 mm \u00d7 10 mm. Lattice structure elements are shown in Figure 2. The lattice cores were fabricated using a Stereo lithography (SLA) 3D printer (Nanotech RS6000) with a low-viscosity stereo lithography resin (Somos\u00ae GP plus 14122). The face sheet is aluminum ESAET 2020 Journal of Physics: Conference Series 1549 (2020) 032121 IOP Publishing doi:10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001309_acc45564.2020.9147355-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001309_acc45564.2020.9147355-Figure3-1.png",
        "caption": "Figure 3. The forces exerted on the follower.",
        "texts": [
            " Without loss of generality, following assumption is made. Assumption 1. External forces exerted on the leader and follower, except the tension of cable, are considered as disturbances, and thereby, not modeled (e.g., drag, wind gust). Assumption 2. Both ends of the cable connecting the leader and the payload can be considered as hinges, which can not resist moments at the points. Therefore, the two forces on both ends of the cable are collinear (i.e., two-force member). The free body diagram of the follower is depicted in Fig.3, where the tension of cable (i.e., FF ) that connects to the follower is the only external force based on Assumption 1. However, disturbances caused by the unmodeled disturbances or noises are tackled by the controller developed in Section IV. The dynamics of the follower is used by the UKF for state estimation and can be described as follows. p\u0307F = vF (1) v\u0307F = 1 mF RF (  0 0 TF \u2212 kvBF )\u2212  0 0 g + 1 mF FF (2) where pF\u2208 R3, mF \u2208 R are the position, and mass of the follower, respectively, and RF \u2208 SO(3) is the rotation matrix from body-fixed frame on the follower to inertial frame, TF\u2208 R is the net thrust, k \u2208 R is the air drag coefficient, vBF \u2208 R3 is the linear velocity expressed in the body-fixed frame on the follower, g \u2208 R is the gravity and FF \u2208 R3 is the external force to be estimated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000016_chicc.2019.8865860-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000016_chicc.2019.8865860-Figure1-1.png",
        "caption": "Fig. 1: Schematic representation of the Stewart platform.",
        "texts": [
            " In order to solve the problems of synthetic position and attitude errors caused by theoretical model errors, the full closed-loop control strategy is adopted in this paper. To the designed controller, the online tuning method to identify the controller parameters is adopted based on RBF neural network. More specifically, RBF neural network is used for system identification to get Jacobian information, and the PID parameters are tuned dynamically with gradient method based RBFNN. Finally, the simulation model is built by Matlab-Simulink and the experiments are designed to verify the control performance. As shown in Fig. 1, base platform B(X,Y, Z) is based on the plane where the center of the six lower hinge points is located, and the plane of the upper is the moving one P (Xp, Yp, Zp). The vectors connecting the upper and lower corresponding order hinge points are the leg vectors. To acquire the rotation matrix of the moving platform, the rotation Angle (\u03b1) around the Z-axis is called Roll, the rotation Angle (\u03b2) around the Y-axis is called Pitch, and the rotation Angle (\u03b3) around the X-axis is called Yaw. The rotation matrix of the moving platform relative to the base platform is as follows: RB p = Rz(\u03b3)Ry(\u03b2)Rx(\u03b1) (1) RB p = \u23a1 \u23a3c\u03b2c\u03b3 s\u03b1s\u03b2c\u03b3 \u2212 c\u03b1s\u03b3 c\u03b1s\u03b2c\u03b3 + s\u03b1s\u03b3 c\u03b2s\u03b3 s\u03b1s\u03b2c\u03b3 + c\u03b1s\u03b3 c\u03b1s\u03b2c\u03b3 \u2212 s\u03b1s\u03b3 \u2212s\u03b2 s\u03b1c\u03b2 c\u03b1c\u03b2 \u23a4 \u23a6 (2) To calculate the lengths of six legs, the closed-loop vector method shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000133_s11071-019-05325-7-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000133_s11071-019-05325-7-Figure8-1.png",
        "caption": "Fig. 8 Projection of the trajectory onto the plane (v1, \u03c9r) and the trajectory of the point of contact for the fixed parameters (20), a = 0.4, c = 0.15 and the initial conditions v1(0) = 2.36, \u03c9s(0) = 1, \u03c9r(0) = 1, \u03d5(0) = 1.86, \u03c8(0) = 0, x(0) = 0, y(0) = 0",
        "texts": [
            " To do so, we choose a parameter of the family, for example, \u03a9 = 1, and fix part of the system parameters Is = 3, Ir = 2, ms = 3, mr = 1, s = 3, b = 1. (20) On the plane of the remaining parameters, (a, c), we will denote regions in different colors depending on the stability of the equilibrium points \u03a3 \u2032 2 and \u03a3 \u2032 3 (see Fig. 7): \u2013 red if both equilibrium points\u03a3 \u2032 2 and\u03a3 \u2032 3 are stable; \u2013 dark blue if both equilibrium points \u03a3 \u2032 2 and \u03a3 \u2032 3 are unstable; \u2013 light blue if \u03a3 \u2032 2 is stable and \u03a3 \u2032 3 is unstable; \u2013 green if \u03a3 \u2032 2 is unstable and \u03a3 \u2032 3 is stable. Figure 8 shows a trajectory with initial conditions from the neighborhood of \u03a32 for parameters for which this equilibrium point is unstable. We see that in this case the trajectory asymptotically tends to one of the equilibrium points \u03a31. At the equilibrium points \u03a31, \u03a3 \u2032 2 and \u03a3 \u2032 3, there is no rotation of the rotor relative to the platform, and hence, we obtain motions of an actually free sleigh (without a rotor) with some fixed mass distribution. Thus, these equilibrium points of the reduced system are generalizations of the equilibrium points of the free sleigh",
            " The exponents \u039b2, \u039b3 are zero to within the calculation error; one of them corresponds to preservation of the energy integral, and the other, to a shift along the trajectory of the Kaplan\u2013Yorke dimension [32] for the corresponding Poincar\u00e9 map: D = 1 + \u039b1 |\u039b4| \u2248 1.26. Atypical viewof the trajectoryof thepoint of contact in this case is presented in Fig. 10b. We see that this trajectory is described by a noncompact curve. In this paper, the problem of the motion of a sleigh with a free rotor has been discussed. It is shown that three types of motion can be distinguished for an unbalanced sleigh: (1) Asymptotically stable equilibrium points in which there is no rotation of the rotor relative to the platform (see Fig. 8). These motion regimes are generalizations of the motion of a usual sleigh (without a rotor). (2) The rotor undergoes periodic oscillations (see Fig. 9), and the trajectory of the point of contact of the sleigh traces out a quasi-periodic curve on the plane. (3) The rotor undergoes chaotic oscillations (see Fig. 10), and the trajectory of the point of contact of the sleigh traces out an unbounded curve on the plane. It is also shown that for this system there exists an integrable case, with a balanced rotor (a = 0, c = 0), which is examined in detail in Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000240_idaacs.2019.8924321-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000240_idaacs.2019.8924321-Figure1-1.png",
        "caption": "Figure 1. Peculiarities of caterpillar MR: cross section on a draft is made by the hinge 3 (a) and robot\u2019s placement on vertical FS (b).",
        "texts": [
            " MOBILE MULTIPURPOSE CATERPILLAR ROBOT Usually permanent magnets are set on tracks in constructions of caterpillar MRs for moving on ferromagnetic surfaces (FSs) [5, 18]. The advantage of such MRs is the high stability on a flat surface and patency due to the large clamping area to the working FS. But there are low reliability at presence of track skews, energy efficiency, life time and rapidity as disadvantages of such solution. So, the authors consider multipurpose (used at least for two technological operations) caterpillar MR (Fig. 1, a, b), in which the main clamping magnet 1 is supported by spherical joint 3 on the frame 4 bottom between right and left tracks 5 with providing the clearance relative to the FS 2 (Fig. 1, a). The presence of the appropriate clearance between the separate main permanent magnets (two may be used or more) and the FS can decrease consumption of energy and increase the passability as well as moving speed of MR. The main properties and the principal diagram of the multipurpose caterpillar MR\u2019prototype are described in [18] in detail, its modernized version as an experimental model is shown in Fig. 1, b. III. STRUCTURE AND SYNTHESIS METHOD OF THE COMPLEX NEURAL CONTROLLER OF THE MR\u2019S SPATIAL MOTION The functional structure of the two-channels automatic control system of the MR\u2019s spatial motion on the basis of complex neural controller is presented in the Fig. 2. The following designations are used: US and U are the setting signals corresponding to the given values of the MR\u2019s speed V RG and angular coordinate RG, which come from the upper control level; S and A are the error signals in speed and angle channels accordingly, which arrive to the motion neural controller (MNC); UC 1 and UC 2 are the voltages supplied to the MR\u2019s model for the each thyristor converter, which supply MR\u2019s drive motors of rotating tracks; UAS and USS are the feedback signals from the speed (SS) and angle (AS) sensors; R and V R are the MR\u2019s current angular coordinate (robot\u2019s course) and linear speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000462_s10556-020-00699-7-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000462_s10556-020-00699-7-Figure3-1.png",
        "caption": "Fig. 3. Improved construction of standard slider bearing for effective delivery and removal of oil: (a) \u2014 schematic diagram of bearing of high-pressure compressor of UKSI-16/500 plant for cycling process with oil scrapers; (b) \u2014 schematic diagram of plant of TRIZ\u00ae oil scrapers in standard bearing; (c) \u2014 TRIZ\u00ae oil scrapers for standard bearing.",
        "texts": [
            " In order to dampen the vibrations, traditional five-pad bearings were replaced by three-pad bearings with supports resting on a self-generated oil layer, and despite this step, the temperature of the pads increased to 100 \u00baC in the course of operation of the rotor, which still did not fall outside the operating regime (Table 3). In order to reduce the temperature in the bearing, specialists at the TRIZ firm made the decision to increase the diametral clearance to the level of buoyancy of the pads in a self-generated damping oil layer and also used oil scrapers for high-efficiency heat extraction (Fig. 3). In order to implement this design, the distributor groove near the inlet edge of the pad was produced with a slotted channel directed from the groove to the face of the pad opposite the direction of rotation of the shaft and recesses were created in the outlet edge in which an oil scraper was installed for effective removal of the hot lubricant from the spindle (i.e., to reduce the temperature of the pads). The scraper was constructed from wearresistant, anti-scuff electrically conducting material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure15-1.png",
        "caption": "Fig. 15. Schematic configuration of (a) the proposed HEBFM (b) the 24- slot/10-pole wound field machine.",
        "texts": [
            " (IDC1, IDC2) amplitude (FEM) phase angle (FEM) amplitude (analytical) phase angle (analytical) (5A, 5A) 8.44V 0.9deg 8.59V 0.88deg (5A, 0A) 6.64V 7.8deg 6.65V 7.5deg (5A, -5A) 5.01V 25.1deg 5.10V 24.3deg (0A, 5A) 6.65V -7.2deg 6.65V -6.9deg (0A, 0A) 4.58V 0deg 4.63V 0deg (0A, -5A) 2.61V 16.2deg 2.71V 15.8deg (-5A, 5A) 5.03V -23.3deg 5.10V -22.3deg (-5A, 0A) 2.65V -17.6deg 2.75V -16.9deg (-5A, -5A) 0.95V 0deg 0.97V 0deg In this part, the performance of the proposed HEBFM and the 24-slot/10-pole wound field machine (WFM) are studied and compared. Configuration of two machines are shown in Fig. 15. Both of the machine are designed with the same outer diameter, axial length and air gap length. The material employed in each part of both tested electric machines are the same. In addition, both electric machines are optimized based on the multiobjective differential evolution couple with FEM, which will be elaborated in the next section. Both machines are operated at the temperature 25 \u00baC, rated speed 300 rpm, filed current 5A and armature current on the amplitude of 6A. Torque, torque ripple and efficiency are measured and calculated, where the results are listed in Table VI"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure8-1.png",
        "caption": "Fig. 8: calculated a) and measured b) breathing mode shape",
        "texts": [
            " In addition, the end shields and the flange were considered by an elastic support boundary condition indicated by the blue color in the screw threads of the housing. Moreover, this boundary condition allows torsional mode shapes to occur in the simulation. The experimental setup is shown in Fig. 7. Triaxial accelerometers were placed on the outer housing surface of the machine. At 42points (3 rings with 14 points), the machine was excited using a B&K 8206 impact hammer. The software B&K Connect was used to evaluate the stable mode shapes and eigenfrequencies. In Fig. 8 and Fig. 9, the comparison between the measurement and the simulation are displayed. It has to be noted that the measurement was executed with the motor freely mounted by flexible straps (Fig. 7). However, the FE-analysis includes an elastic boundary condition. Likewise, the experimental analysis was performed on the assembled machine including the end shields as well as the rotor. Because of this additional mass, only three mode shapes are identified confidently. Nonetheless, the simulation shows good accordance with the measured eigenfrequencies, as indicated in table III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001153_01691864.2020.1782260-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001153_01691864.2020.1782260-Figure14-1.png",
        "caption": "Figure 14. Demonstration key poses inserted (left in each subfigure) and the collision-free motion paths (right in each subfigure) generated with (a) Policy 1 and (b) Policy 2 of peg-rot. Both use the planning result in first row under each policy.",
        "texts": [
            " Table 3 shows the results of motion planning of peg-rot. Different from the two tasks above, Policy 1 and Policy 2 made use of similar number of demonstration poses and their best performances cost almost the same single planning time and total time. However, Policy 2 was more stable while Policy 1 sometimes cost longer single planning time and total time. Demonstration key poses inserted and the eventual obtained collision-free motion paths of peg-rot generated with Policy 1 and Policy 2 are shown in Figure 14. Table 2. Motion planning result of tenon insertion. Selection policy Inserted pose number Single planning time (s) Total time (s) Policy 1 3 0.473 9.949 3 0.472 10.019 3 0.429 9.719 3 0.395 9.413 3 0.457 9.652 Policy 2 2 0.425 4.875 2 0.449 5.017 2 0.425 4.837 2 0.473 4.947 2 0.388 5.120 To summarize, Policy 2 generallymade use of fewer poses than Policy 1, thus took less time to complete the motion planning. The explanation is as below: The planner kept inserting the new key poses until the solution path was found",
            " Both use the planning result in first row under each policy. Table 3. Motion planning result of peg-rot. Selection policy Inserted pose number Single planning time (s) Total time (s) Policy 1 1 7.509 7.514 1 0.442 0.447 1 0.467 0.471 1 0.427 0.432 1 4.219 4.219 Policy 2 1 0.497 0.502 1 0.511 0.516 1 0.541 0.547 1 0.416 0.420 1 0.505 0.509 right figure in Figure 15(b)). Admittedly, this feature of Policy 2 is only reflected in specific tasks where there is an apparently optimal path among many latent solutions. Taking a counterexample, in Figure 14, the two paths did not show much difference because all the latent solution paths were very similar to each other. Finally, comparing Figure 15(a-b), it is obviously that the proposed method compensated the pose errors in captured human demonstration, avoiding the interference between objects. To validate the effectiveness of the proposedmotion planning through the demonstration method, actual robot experiments have been executed on the Nextage OPEN humanoid robot. The robot control was realized through the Robot Operating System (ROS) APIs of Nextage OPEN2 in PyManipulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002170_s00170-020-06399-z-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002170_s00170-020-06399-z-Figure5-1.png",
        "caption": "Fig. 5 Illustration of Goldak\u2019s et al. double ellipsoid heat source model (indicating scan motion in \u201cY\u201d direction)",
        "texts": [
            " The SLM modeling presented here predicts the distortion distribution according to the transient thermal history, which depends on many factors including the build geometry, scan speed, scan pattern, powder layer thickness, laser power, temperature-dependent material properties, ambient conditions, etc. A moving heat source model is applied here to the SLM process based on Goldak et al.\u2019s double ellipsoid Gaussian heat source [41, 42], which was originally introduced in 1984 for welding simulations. The double ellipsoid model, which describes thermal power density distribution in terms of a volumetric heat flux, is depicted in Fig. 5 and quantitatively described by Eq. 7: qf (t) = 6 \u221a 3ff P\u03b1 abcf \u03c0 \u221a \u03c0 e \u22123 x2 a2 e \u22123 z2 b2 e \u22123 (y+vt)2 cf 2 , y \u2265 yi qr (t) = 6 \u221a 3frP\u03b1 abcr\u03c0 \u221a \u03c0 e \u22123 x2 a2 e \u22123 z2 b2 e \u22123 (y+vt)2 cr 2 , y < yi for t \u2265 0. (7) In Eq. 7, q is the volumetric heat flux, P is the incident laser power, yi is the instantaneous Y ordinate (assuming the scan proceeds in the Y direction), and ff and fr are fractions of the heat distribution within the front and rear octants, respectively. A term \u03b7, representing absorptivity of the powder particles based on mean particle size, is also included, [43, 44]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001859_j.promfg.2020.10.106-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001859_j.promfg.2020.10.106-Figure2-1.png",
        "caption": "Fig. 2. Powder hopper/doser subsystem: (a) 3D model, (b) photo.",
        "texts": [
            " Then, a recoating cylinder (roller) in combination with a doctor blade is responsible for the creation of a smooth, homogeneous and evenly compressed powder layer. The sieve/doctor blade/roller assembly moves horizontally and parallel to the powder bed compartment via the x-axis motor, which is not depicted in drawings for simplicity. Note that all custom designed parts were manufactured inhouse on a HaasTM TM-1 machining center. The powder hopper/doser subsystem consists of the powder tank in which powder is stored and the doser drum, which is a cylinder with a helical pattern of blind holes of fixed volume that is rotated by a stepper motor, see Fig. 2. The powder tank can be sealed and equipped with humidity and temperature sensors in order to measure the environmental factors that actively affect the powder quality. Panagiotis Avrampos et al. / Procedia Manufacturing 51 (2020) 755\u2013762 757 2 Author name /Procedia Manufacturing 00 (2019) 000\u2013000 emptied and refilled regularly, otherwise the conditions of the raw powder can be considered neither stable nor known. Fig. 1. Powder delivery piston with recoater (doctor) blade in SLS machine. Commercial PDS come with a roller and/or a doctor blade, as means of applying an evenly spread powder layer",
            " The sieve/doctor blade/roller assembly moves horizontally and parallel to the powder bed compartment via the x-axis motor, which is not depicted in drawings for simplicity. Note that all custom designed parts were manufactured in- house on a HaasTM TM-1 machining center. 2.1. Powder hopper/doser The powder hopper/doser subsystem consists of the powder tank in which powder is stored and the doser drum, which is a cylinder with a helical pattern of blind holes of fixed volume that is rotated by a stepper motor, see Fig. 2. a b Fig. 2. Powder hopper/doser subsystem: (a) 3D model, (b) photo. The powder tank can be sealed and equipped with humidity and temperature sensors in order to measure the environmental factors that actively affect the powder quality. Author name / Procedia Manufacturing 00 (2019) 000\u2013000 3 According to these measurements, it is possible to fit a dehumidifier or just control the laser beam of the SLS machine in such a way as to counter the effects of the humidity in the powder. The tank is fitted with a powder level gauge to monitor when it needs to be refilled"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure11-1.png",
        "caption": "Fig. 11 The planar double-folded metamorphic mechanism",
        "texts": [
            " According to the above steps, the position, velocity, acceleration, the driving force/torque of active parts and the constrained force/torque of metamorphic joints can be obtained at time tn by iteration. 1 3 The iteration flowchart is shown in Fig.\u00a010. Taking the planar double-folded metamorphic mechanism (Wang and Dai 2007) as an example, the types and constraints of metamorphic joints based on the kinematic characteristics of metamorphic joints are given. According to the working task, the planar double-folded metamorphic mechanism (as shown in Fig.\u00a011) must complete three working requirements: horizontal folding, vertical folding and reset. (1) Horizontal folding, as shown in Fig.\u00a011a, horizontally pushing the left side of the single-layer cardboard and rotating the second side of the left side along the second crease to vertical position. (2) Vertical folding, as shown in Fig.\u00a011b, completing the rotation of the first surface around the first crease and coinciding with the second surface. (3) Reset, as shown in Fig.\u00a011c, the mechanism returns to its initial position. Figure\u00a012 shows the schematic diagram of the planar double-folded metamorphic mechanism and its composition principle. According to the above requirements, force metamorphism is adopted at kinematic joint D, that is, when the mechanism is in configuration 1, the spring force is set at kinematic joint D, so that the relative moving resistance between the components 3 and 4 is larger than the motion resistance of the slider, and it keeps the components 3 and 4 relatively static"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001619_s11668-020-01014-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001619_s11668-020-01014-5-Figure1-1.png",
        "caption": "Fig. 1 (a) Schematic diagram of conveyors. (b) Schematic diagram of failed pulley assembly (c). Overview of conveyer shows location of drive side bearing. (d) Closer view of failure location and pulley position",
        "texts": [
            " This flaking progressively increases in extent and eventually makes the bearing unserviceable. Such an incident happened in a conveyor system in an integrated steel plant. Bearings of both drive side and non drive side of weigh feeder pulley were found failed after 4 years of service. There were a total of 48 bearings of which three bearings were found damaged. The conveyors are used to transfer coke breeze feed to the stacking yard. A pulley is used to provide tension in the system. A schematic diagram and the location of failure are shown in Fig. 1. The failed components were collected from the plant for investigation. The samples were cleaned for visual examination. Samples for microstructural evaluation were prepared from the inner race, outer race and balls. These were mounted and polished by conventional U. Pal (&) P. Palit P. Gokan Central Laboratories, Scientific Services Division, Tata Steel Limited, Kalinganagar, India e-mail: urbi.pal@tatasteel.com S. Kanrar Mechanical Design, Tata Steel Limited, Kalinganagar, India metallographic techniques to obtain a scratch-free surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001727_s0263574720000843-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001727_s0263574720000843-Figure2-1.png",
        "caption": "Fig. 2. Kinematics diagram.",
        "texts": [
            " In the previous design of this robot, the optimum configurations for the six kinematic chains are determined as follows: the distribution of all the spherical joints is on a circle with radius rm theoretically. The center of this circle is point P. Three of the S joints (denoted by M1 M2 M3) are distributed symmetrically on this circle as well as the other three S joints (denoted by M4 M5 M6). And each central angle of the adjacent S joints (denoted by \u2220M1 P M4, \u2220M2 P M5, \u2220M3 P M6) is 30\u25e6, as shown in Fig. 2. Theoretically, in the initial position of the robot, for each kinematic chain, the linkage is collinear with the P joint and is tangent to the circle mentioned above. This configuration can realize singularity-free in the demanded workspace. The dimensions and workspace of this robot will be introduced later. For parallel robots, the inverse kinematics is easier to obtain than forward kinematics, so we will take advantages of the inverse kinematics to deduce the calibration model. The next is a concise inverse kinematics description. The moving coordinate frame P{x \u2032, y\u2032, z\u2032} and the fixed coordinate frame O{x, y, z} are shown in Fig. 2. The origin of the moving frame is point P and the origin of the fixed frame is point O. In the initial position of the robot, point P theoretically coincides with point O. The illustration for the inverse kinematics is based on any one of the kinematic chains of the robot. We use the subscript i to denote the ith kinematic chain. Here, i is from 1 to 6. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000843 Downloaded from https://www.cambridge.org/core. San Francisco State University, on 12 Nov 2020 at 14:28:48, subject to the Cambridge Core terms of use, available at As shown in Fig. 2, point Mi is the theoretical center of the ith S joint. Point Di is the theoretical rotation center of the ith U joint. The initial position of point Di is marked by point Gi . In the fixed frame O{x, y, z}, the position vector of the point Gi is represented by gi . Point Ni is the projection point of Mi onto the direction vector of the ith P joint. In the fixed frame O{x, y, z}, the vector \u2212\u2212\u2192 P Mi is represented by mi , while it is denoted by mi \u2032 in the moving frame P{x \u2032, y\u2032, z\u2032}. The vector \u2212\u2212\u2212\u2192 Di Mi is denoted by li with respect to the fixed frame O{x, y, z}",
            " The vector \u03b8 represents the angle coordinates of the https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000843 Downloaded from https://www.cambridge.org/core. San Francisco State University, on 12 Nov 2020 at 14:28:48, subject to the Cambridge Core terms of use, available at moving platform. Unit vector ni represents the direction of the ith linkage. So, there are the following equations: \u03b8 = [ \u03b1, \u03b2, \u03b3 ]T li = Li ni , (i = 1, 2, . . . , 6) (7) From the kinematics diagram in Fig. 2, we can get p + Rmi \u2032 = gi + qi ki + li = gi + qi ki + Li ni , (i = 1, 2, . . . , 6) (8) Taking the derivative of both sides of Eq. (8), there is \u03b4p + \u03b4Rmi \u2032 + R\u03b4mi \u2032 = \u03b4gi + qi\u03b4ki + \u03b4Li ni + Li\u03b4ni , (i = 1, 2, . . . , 6) (9) where \u03b4 is the symbol of differential operator. According to the mathematical definition of the rotation matrix R, there is the following equation: \u03b4R = \u23a1 \u23a2\u23a2\u23a3 0 \u2212\u03b4\u03b3 \u03b4\u03b2 \u03b4\u03b3 0 \u2212\u03b4\u03b1 \u2212\u03b4\u03b2 \u03b4\u03b1 0 \u23a4 \u23a5\u23a5\u23a6 R = [\u03b4\u03b8\u00d7] R (10) where [\u03b4\u03b8\u00d7] represents the skew-symmetric matrix in this equation",
            " For Group 1, it is a basic mathematical problem for linear fitting. And the corresponding toolbox is embedded in the supporting software of most measuring apparatus. The information of the constructed space line can be displayed as soon as the sample points are measured. That is quite convenient and efficient. For Group 2, the rest parameters which contain a reduced number of unknowns need to be identified by certain error model. Here, we present an error model by taking advantage of the inverse kinematics of the robot. According to the kinematics diagram in Fig. 2, there is the following relation: Li 2 = \u2223\u2223p + Rmi \u2032 \u2212 gi \u2212 qi ki \u2223\u2223 , (i = 1, 2, . . . , 6) (15) After measuring an actual pose of the robot, the components of p and R can be determined and then be written as follows: \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 p = [px , py, pz]T R = \u23a1 \u23a2\u23a3 r11 r12 r13 r21 r22 r23 r31 r32 r33 \u23a4 \u23a5\u23a6 , (i = 1, 2, . . . , 6) (16) where the nine components of the rotation matrix R are calculated according to the roll-pitch-yaw angles. The coordinates of the vectors m\u2032 i , gi and ki can be denoted by:\u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 mi \u2032 = (mxi \u2032, myi \u2032, mzi \u2032)T gi = (gxi , gyi , gzi ) T ki = (kxi , kyi , kzi ) T , (i = 1, 2, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001707_0309324720958257-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001707_0309324720958257-Figure4-1.png",
        "caption": "Figure 4. Schematic diagram of contact between braided wire rope and drum rope groove.",
        "texts": [
            " According to equations (1)\u2013(4), the stress and strain of the core wire and side wires of the inner and outer layers under the tensional load of the straight strand can be calculated. That is, the mechanical response of the rope under the tensile load of the strand is obtained. Mechanical response of braided wire rope strand under torsional load During the process, the wire rope is subjected to an axial torsion due to the friction of the drum. Therefore, it is necessary to analyze the mechanical response of the strand under the torsional load. First, the torque of the strand is analyzed. The contact state of the braided wire rope and reel is shown in Figure 4. It can be seen from Figure 4 that the contact between the drum and braided wire rope is the same as the contact between the drum and rope strand. The number of rope strands in contact with drum groove is different at different positions, but within a pitch, each rope strand has the opportunity to contact with the rope groove. Therefore, the external torque of the rope strand is the same as the torque of the drum friction to the rope strand, and the torque of the wire rope is the sum of the torque of each rope strand. The specific stress conditions are shown in Figure 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001528_0954407020951318-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001528_0954407020951318-Figure7-1.png",
        "caption": "Figure 7. DMF test bench.",
        "texts": [
            " Finally, the total friction torque acted on arc-spring can be obtained: Tf =Tf1 +Tf2 \u00f012\u00de In addition to the friction torque of the arc-spring, there is also basic friction damping torquegenerated by the friction disc in the twisting process of the DMF. The basic friction damping torque is mainly used to reduce the resonance peak. The expression of the friction torque is: Tfir =M0 \u00f013\u00de In summary, the nonlinear torsion model of DMF can be obtained: TDMF= k(u)+ (Tf +Tfir)sign( _u) \u00f014\u00de where, _u is the twist angle velocity between the first mass and the second mass. The test sample is a multi-stage stiffness DMF. It is verified by experiments that the DMF model is correct.The bench test is shown in Figure 7. The second mass of the DMF is bolted to the mounting bracket. The first mass is connected with the torsion actuator, and the test bench is equipped with an angle sensor and a torque sensor. The actuator is slowly loaded, and its angular displacement loading curve is as shown in Figure 8. The data acquisition instrument is used to collect angular displacement and torque data of DMF, and its torsional characteristic curve is obtained through data processing, as shown in Figure 9. The parameters of the DMF are obtained through the test, and the comparison results with the design values are shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002093_telecom50385.2020.9299536-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002093_telecom50385.2020.9299536-Figure1-1.png",
        "caption": "Fig. 1. Discontinuity points of the robot distance sensor reading as radar.",
        "texts": [
            ") obstacle detection such as Long-Range Ultrasonic Time-of-Flight Range Sensor capable to sens long obstacle boundaries; \u2022 9-DOF Inertial Measurement Unit with 3-axis accelerometer, 3-axis magnetometer, and 3-axis gyroscope; \u2022 LiDAR and RGB Camera with for precise volumetric measurements of objects; \u2022 LED IR light source with IR (NoIR) camera at the bottom; Our pathfinder strategy adds a few steps taken from the colony intelligence techniques on top of mentioned in section I. Definitions of the terms and equations that are used in the algorithm are as follows: X : is the current position of the robot. L angle of interest - direction to the goal and can be set grammatically Qgoal: is the goal position. Qtmp: is the temporary goal position for the advance avoiding of a recognized obstacle. O: is the i-th discontinuity point as it is shown in figure 1. The discontinuity points are the locations where the sensor information suggest that a continuity interval has ended and another interval has started D: is the estimated distance from the current position to the goal through the local tangent point with the heuristic function D(x,Qgoal) = D(x,Oi) +D(Oi,Qgoal) (1) Authorized licensed use limited to: Carleton University. Downloaded on June 05,2021 at 14:59:56 UTC from IEEE Xplore. Restrictions apply. Dreach : is the output of the heuristic function taking into consideration the best local tangent point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001431_tec.2020.3017077-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001431_tec.2020.3017077-Figure2-1.png",
        "caption": "Fig. 2. Stator and rotor of three phase asynchronous motor YHPE 300-4",
        "texts": [
            " The disk motor is composed of a disk-shaped stator and a disk-shaped rotor. The axial gap between the two disks is the working air-gap of the motor, as shown in Fig. 1. The cores of stator and rotor are formed by rolling the silicon steel strip (DW310) of which the slots are punched. The conventional three-phase asynchronous disk motor of YHPE 300-4 is selected as the cryogenic driving motor. The main design parameters are listed in Table I. The photograph of the stator and rotor for the disk motor are shown in Fig. 2. To facilitate the automatic rolling process, semi-closed parallel-sided flat-bottomed and circular-bottomed slots are adopted in the stator and rotor core of disk motor respectively. The main structural dimensions of the disk motor are illustrated in Table II. When the stator winding of disk motor is powered, there are axial and radial electromagnetic forces between the stator and the rotor, besides the useful electromagnetic torque. This paper mainly focuses on the electromagnetic force behaviors of the disk motor using for cryogenic submersible pump"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002488_ipemc-ecceasia48364.2020.9367925-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002488_ipemc-ecceasia48364.2020.9367925-Figure1-1.png",
        "caption": "Fig. 1. Structure of machines. (a) TPM. (b) UT-FTPM. (c) STO-FTPM.",
        "texts": [
            " And it must be pointed out that the fault-tolerant machines mentioned above improve the fault-tolerant capability at the expense of large torque loss. In this paper, the fault-tolerant tooth and the stator tooth offset are proposed to achieve the better fault-tolerant capability. This paper introduces the structure of TPM, UT-FTPM and STO-FTPM. Their electromagnetic properties, including output torque, selfinductance and mutual inductance, will be compared and analyzed. II. STRUCTURE OF MACHINES Fig. 1(a) shows the structure of the TPM. Two coils of different phases are placed in the same slot, which leads to strong electromagnetic coupling. Fig. 1(b) shows the structure of UT-FTPM. Each coil is wound around per teeth and each slot only contains a single phase. Fault-tolerant teeth realizes the physical isolation, magnetic isolation and thermal isolation between phases. Fault-tolerant tooth only provides a path for the magnetic flux loop, so the thickness of the tooth body and the width of tooth tips are narrower than those of the armature tooth, which reduces the weight of the machine and the amount of iron. However, the output torque of this machine will be reduced because the stator slot area is not sufficiently utilized. Fig. 1(c) shows the structure of the STO-FTPM. The slot area between armature tooth increases and the slot area between armature teeth with fault-tolerant teeth decreased through the method of stator tooth offset. Therefore, the slot area is sufficiently utilized and the output torque is improved compared with UT-FTPM. Fig. 2(a) shows the partial expansion diagram of the TPM. \u03b8_OT is the width of teeth tips and D_T is the depth of teeth tips. Fig. 2(b)is the partial expansion diagram of the UTFTPM. The fault-tolerant tooth only provides a path for the magnetic flux loop and the tips of the fault-tolerant tooth passes fewer magnetic force lines, so the thickness of the faulttolerant tooth W2 and the width of tips \u03b8_FT are both smaller than the thickness of the armature tooth W1 and the width of tips \u03b8_AT"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000170_ecce.2019.8913254-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000170_ecce.2019.8913254-Figure8-1.png",
        "caption": "Fig. 8. Case 2- Study Domain",
        "texts": [
            "8 mm (57 inches) The amplitude of the forces on the bottom bar would be using (4): = 216 N/m Using kf factor provided on Table 1 and the core length provided above, the amplitude of the forces on the slot content can be calculated. The amplitudes of the force are provided on Table 5. Similarly to case 1, a 2D time stepped electromagnetic finite element analysis was performed to investigate the contribution of the stator core teeth saturation, the shape of the rotor pole tip, and the non-uniform damper bar currents on the radial and tangential forces. The software Flux 2D from Altair was used for this investigation. First the study domain was modeled and consisted of a section with 7 poles (Fig. 8). The side of the model are subjected to periodicity boundary condition with anticyclic conditions. A relatively fine mesh was used in particular on the damper bars and slot wedges. The total number of nodes exceeded 365,000 using with 2nd order finite elements. The increased number of nodes in relation to case 1 was due to the increase of the number of slots in the study domain. Fig. 9 shows details of the meshing. Similarly to case 1, external circuit connections were used on the stator winding, field winding and damper winding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002194_s12206-020-1201-5-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002194_s12206-020-1201-5-Figure8-1.png",
        "caption": "Fig. 8. Gears after experiment.",
        "texts": [
            " Sensors 1# and 2# are installed at the input shaft end, sensor 1# is horizontal, sensor 2# is axial direction, sensor 3# is installed at the top of the gearbox, sensor 4# is installed at the output shaft end and is axial direction, and all sensors are installed near the ring gear. In the life-cycle degradation experiment process, the input shaft speed of planetary gearbox is about 1000 rpm, and the load current of the magnetic powder brake is 1 A (about 340 Nm). The sampling frequency of the vibration signal is 20 kHz, each time signal acquisition lasts for 12 s, and it is collected once every 5 minutes. The life-cycle degradation experiment lasts a total of 1003 h. The wear degree of planetary gearbox main gears after the experiment is shown in Fig. 8. The sun gear tooth surface wears the most, the ring gear's tooth surface wears the least, both sides of planet gear tooth surfaces wear, and the side that meshes with the ring gear wears more severely than the other side. In order to reduce the amount of calculation and control the length of the article, this paper only analyzes the degradation process data of sensors 1# and 3#, and only takes one data module every 1 hour. Root mean square (RMS) [29] and energy [30] are conventional performance degradation features and their calculation methods are shown in Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure3-1.png",
        "caption": "Fig. 3. Configurations of 2UPR-PRU PKM, after rotation around: (a) y-axis; (b) u-axis; (c) y- and u-axes.",
        "texts": [
            " Based on Eqs. (2) and (3), the overall constraint wrenches can be expressed as: ( ) ( ) ( ) 3 3 1 2 3 c 0 s ; 0 0 0 0 1 0; tan 0 0 0 0 0; s 0 c \u23a7 = \u2212 \u23aa \u23aa = \u2212 \u2212\u23a8 \u23aa =\u23aa\u23a9 C C A A C x z \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ $ . (4) The overall motion of the proposed mechanism can thus be expressed as: ( ) ( ) ( )3 3 1 2 3 0 0 0; s 0 c 0 1 0; 0 0 0 c 0 s ; 0 c s 0 \u23a7 = \u23aa\u23aa =\u23a8 \u23aa = \u2212 +\u23aa\u23a9 pm pm pm A Az x \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 $ $ $ . (5) Therefore, the 2UPR-PRU PKM has three outputs: two rota- tions around the y- and u-axes ( \u03b2 and \u03b3 ), as shown in Fig. 3, and a translation, which indicates its possible use for applications with orientation requirements. 2.3 Inverse kinematics Inverse kinematics of the 2UPR-PRU PKM can be established using the mapping relationship between the actuated variables T 1 2 3= \u23a1 \u23a4\u23a3 \u23a6q q qq and the output parameters T = \u23a1 \u23a4\u23a3 \u23a6oz\u03b2 \u03b3\u03b7 of the moving platform. Here, oz represents the distance between points O and o along the z-axis and is defined as operational height. The rotation matrix between the o-uvw and O-xyz can be written as: c s s s c 0 c s s c s c c \u23a1 \u23a4 \u23a2 \u23a5= \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 O o \u03b2 \u03b2 \u03b3 \u03b2 \u03b3 \u03b3 \u03b3 \u03b2 \u03b2 \u03b3 \u03b2 \u03b3 R "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000189_ecce.2019.8912698-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000189_ecce.2019.8912698-Figure2-1.png",
        "caption": "Fig. 2. Computational Fluid Dynamic (CFD) mesh of the model",
        "texts": [
            " It is given as a function of \u03bd the kinematic viscosity and u the friction velocity near the wall by the expression [15]: r+ = u.\u0394r / \u03bd (6) This quantity is determined only after performing the first simulation and it is adapted iteratively. To characterize properly the flow structure in the vicinity of both rotor and stator walls. It is important to notice that r+ is taken strictly less than 1 (r+<1). A structured quadrilateral mesh is used for all configurations discussed in this article. To resolve correctly the field near the walls, a finer mesh layers with a growing factor of 1.2 was applied (Fig. 2). D. Setting Up Conservation equations are solved numerically using the LES approach due to its ability to resolve smaller structures compared to other approaches. The WALE (wall-adapting local eddy-viscosity) subgrid scale model was chosen. Simulations were run with transient pressure-based solver. No slip conditions are applied on rotor and stator walls. The temperature is supposed constant and uniform on rotor and stator surfaces. Periodicity conditions are imposed in both axial and circumferential directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001969_5.0028365-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001969_5.0028365-Figure1-1.png",
        "caption": "FIG. 1. 3D model of the AACMM and its relationship with the adjacent coordinate frame: (a) the 3D model and (b) the initial position of the kinematic model.",
        "texts": [
            " The rest of this paper is organized as follows: the development of the kinematic model and the error model of the AACMM are shown in Sec. II. Section III presents the detailed procedure of the BP neural network optimized by the MEA (BP-MEA). The simulation and experiment for kinematic error calibration and non-kinematic error compensation of the AACMM are shown in Sec. IV. The discussions and conclusions are drawn in Sec. V. The kinematic model of the AACMM is modeled by D\u2013H convention.3,24 The 3D model of the AACMM and its kinematic chain from the base coordinate system (OX0Y0Z0) to the measuring probe (OXpYpZp) are shown in Fig. 1. Suppose that a given set of six joint angles for the AACMM is \u03c9 = [\u03b81, \u03b82, . . . , \u03b86] T . Any spatial coordinate of the measuring probe with respect to the base coordinate system can be formulated as follows: P(\u03c9) = 6 \u220f i=1 Ai\u22121 i A6 p(i = 1, 2, . . . , 6), (1) where A6 p = [0, 0,Zp, 1]T, Zp indicates the length of the measuring probe, and the homogeneous transformation matrix Ai\u22121 i is represented as follows: Ai\u22121 i = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 cos\u03b8i \u2212sin\u03b8icos\u03b1i \u2212sin\u03b8isin\u03b1i licos\u03b8i sin\u03b8i cos\u03b8icos\u03b1i \u2212cos\u03b8isin\u03b1i lisin\u03b8i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , (2) where \u03b8i, \u03b1i, di, and li represent the joint angle, link twist, link offset, and link length of the AACMM, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure10-1.png",
        "caption": "Fig. 10. Demagnetization ratio distribution for DTPSPM machines without eccentricities at 16ms during 3PSC. (a) Type I. (b) Type II. (c) Type III.",
        "texts": [
            " 9 and the demagnetization distributions are 0 20 40 60 80 0 1 2 3 4 5 6 7 8 9 10 11 12 F x ( N ) Harmonic order Fx-before Fx-after -100 -50 0 50 100 0 60 120 180 240 300 360 F y (N ) Rotor position (Mech. Deg.) Fy-before Fy-after Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. shown in Fig. 10. It can be seen that for the type I DTPSPM machine, the demagnetization distribution is rotationally symmetrical whilst for the other two winding configurations, the demagnetization distribution is rotationally asymmetric. This is because each winding group is rotationally asymmetric for the type II and III machines. In addition, since the 3PSC happens only in one winding group, the armature fields during 3PSC are not rotationally symmetrical. As a result, there is no post-demagnetization UMF for the type I DTPSPM machine whilst the post-demagnetization UMF occurs for the other two winding configurations, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002755_tmag.2021.3074935-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002755_tmag.2021.3074935-Figure7-1.png",
        "caption": "Fig. 7. Rated load flux distributions of Models 1\u20134. (a) Model 1: 8/6/2. (b) Model 2: 5/3/2. (c) Model 3: 7/3/2. (d) Model 4:7/6/1.",
        "texts": [
            " Model 4 exhibits the highest fundamental amplitude, thanks to the highest pole ratio (Ge = 7), which are about 1.5\u20132 times higher than the other models. The back EMFs in Models 1\u20133 contain rich harmonics while the back EMFs in Model 4 is more sinusoidal with little harmonics. The fundamental wave of the back EMF is produced by the dc biased current, which will interact with ac current to produce average torque while the high-order harmonics will produce torque ripple. When both the dc and ac components are injected into the windings, average torque will be produced. Fig. 7 shows the flux distributions on rated load (Idc = 13.44 A and Iac = 13.44 A) of models. Similar to no load distribution, the flux distribution of Model 1 is centrosymmetric and that of model 4 is symmetric while for Models 2 and 3, it is asymmetric. Local saturation occurs in the stator tooth in Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 24,2021 at 17:02:44 UTC from IEEE Xplore. Restrictions apply. Models 2 and 3 and the stator yoke in Mode 4 due to asymmetric flux distribution. Fig. 8 exhibits the radial force density of four models at the same rotor position as in Fig. 7. As can be seen, unbalanced magnetic force exists in Models 2\u20134, which might likely produce vibrations and noise. With the rotating exciting fields and rotating armature MMF, the electromagnetic torque ignoring the iron saturation effect can be obtained with the following equation: Tem = 2\u03c0rgleff B A \u221d 2\u03c0rgleff Idc Iac N2 a (8) where A is the armature electrical loading. It can be seen that the torque is proportional to the square of the number of turns, and the biased dc component and the ac component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000547_iros40897.2019.8967922-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000547_iros40897.2019.8967922-Figure8-1.png",
        "caption": "Fig. 8: Resulting trajectory recorded from experiment. The red-dashed line indicates the desired trajectory computed from the proposed path planner, and the blue line shows the actual trajectory of the aerial manipulator. The dark lines of each color denote the trajectories of the body position of aerial manipulator while the light lines represent the position of end effector.",
        "texts": [
            " To control the robotic arm, we implement a simple PID controller based on joint position and velocity measurements. For indoor positioning, we use VICON motion capture system. During the flight, the pose of the aerial manipulator is sent to the ground control station at 100 Hz from the VICON system. The onboard computer receives the pose of the aerial manipulator via wireless connection from the ground control station. The experimental results are reported in Figs. 8\u20139. The resulting trajectory is shown in Fig. 8 and the time histories of the state variables are listed in Fig. 10. In Fig. 8, each solid and light line denotes the trajectory of the body and end effector positions of aerial manipulator. In both figures, the blue lines show the actual trajectories and the red lines represent the desired trajectories computed from the proposed planner. Fig. 9 shows the snapshots of the experiment. With our proposed planner, the aerial manipulator successfully catches the object during the task and finishes the task on the goal position. Incremental sampling-based optimal algorithms have been widely used due to their benefits including the flexibility for formulating the difficult path planning problem, the ability of quickly searching for a feasible path, and the guarantee of a global motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure11-1.png",
        "caption": "Fig. 11. Installation state diagram of the side surface",
        "texts": [
            "10, so that the bottom surface can reach the angle steel where the insulator hanging point is located. The bottom surface of the left and right sides of the device considers the height of the bottom angle steel on the cross arm, which also allows the design and installation of the bottom surface device to ignore the height of the screws that fix the angle steel structure of the tower. Because the angle steel extends to both sides of the cross arm on the side of the device, there is a gap in the area near the upper side of the side surfaces, as shown in Fig.11, which does not affect the angle steel that has been fixed on the transmission line tower. At the same time, let the device be close to the left device as far as possible. Although there is a gap on the side, the angle steel has a space occupied by the gap, and birds cannot enter the protection area of the intermediate phase device through the gap. Due to the existence of the angle structure and the fact that the device is hand-made based on experience, the device accuracy is very low, after installation there are gaps in the connection part of the device, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003203_s42835-021-00864-9-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003203_s42835-021-00864-9-Figure18-1.png",
        "caption": "Fig. 18 Four coil structure a The magnetic density illustration, b topview of the coil",
        "texts": [
            " 15 The variation of Induced voltage of receivers under different sizes (a) The relationship between misalignment distance of C1 and C2 coil and induction voltage in x-axis direction, (b) The relationship between misalignment distance of C1 and C2 coil and induction voltage in y-axis direction, (c) The relationship between misalignment distance of C3 and C4 coil and induction voltage in x-axis direction, (d) The relationship between misalignment distance of C3 and C4 coil and induction voltage in y-axis direction \u25b8 1 3 observed at the distance of 0.72\u00a0m from the center of the shielding structure on the x- axis. Similarly, the distance is 0.63\u00a0m from the center of the coils on y- axis, which are all within the transmitting coil. When the DWPT system operates, four independent coils at the receiving end generates alternating magnetic field, and the direction of the induced electromotive force is therefore alternating. The four coils are connected in series in the experiment. As shown in Fig.\u00a018, the current is connected in the order of 1#-4#-3#-6#-5#-8#, and 2# and 7# are the output power terminals. In the course of the experiment, when the horizontal center of the receiving coil and the transmitting coil are 1 3 perfectly aligned, the system transfers high power. When the receiving coil deviates from the central position, the two receiving coils in the x-axis direction produces voltage differences. The robot position detection and adjustment circuit (PDAC) compares the voltage difference with the preset voltage values stored in the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002338_012093-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002338_012093-Figure3-1.png",
        "caption": "Figure 3. Element and forces acting on each node.",
        "texts": [],
        "surrounding_texts": [
            "External loads may be in the structure at different locations and may not match the nodes of the elements. Despite the various maneuvers we can do, at the distributed load must be equivalent to the joints, because at the joints we have displacements. Equivalence will be done according to potential energy or internal work, so the work of the forces in the scheme is equal to that of the equivalent forces in the joints. A simple method of determining equivalent forces is to find the reactions of the particular element scheme. 10th EASN 2020 IOP Conf. Series: Materials Science and Engineering 1024 (2021) 012093 IOP Publishing doi:10.1088/1757-899X/1024/1/012093 We know that the work of external forces is equal to the work of internal forces (but with a minus sign ahead). External forces are divided into active forces, that do work and reactions that do not. \ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e ,\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f) = \ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e) = \u2212\ud835\udc38\ud835\udc38\ud835\udc5d\ud835\udc5d\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d = \u2212\u2211 \ud835\udc40\ud835\udc402 2\u2219\ud835\udc38\ud835\udc38\ud835\udc38\ud835\udc38 \ud835\udc51\ud835\udc51\ud835\udc51\ud835\udc51 (1) \ud835\udc34\ud835\udc34\ufffd 3 2\ufffd = \ud835\udc34\ud835\udc34(+) \u2212 \ud835\udc34\ud835\udc34(\u2212)(\ud835\udc3c\ud835\udc3c = 3 2 ) (2) We remove the real connection and place imaginary support where the active force act. When we do this, the reactions become active forces and active force became non-active force. In this case reactions \ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e ,\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f) = \ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f) = \u2211 \ud835\udc40\ud835\udc402 2\u2219\ud835\udc38\ud835\udc38\ud835\udc38\ud835\udc38 \ud835\udc51\ud835\udc51\ud835\udc51\ud835\udc51 (3) \u2212\ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f) = \u2212\u2211 \ud835\udc40\ud835\udc402 2\u2219\ud835\udc38\ud835\udc38\ud835\udc38\ud835\udc38 \ud835\udc51\ud835\udc51\ud835\udc51\ud835\udc51 = \ud835\udc34\ud835\udc34(\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f) \u27f9 \ud835\udc39\ud835\udc39\ud835\udc4e\ud835\udc4e\ud835\udc52\ud835\udc52\ud835\udc52\ud835\udc52\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5d\ud835\udc5d\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e = \u2212\ud835\udc39\ud835\udc39\ud835\udc5f\ud835\udc5f\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc5f\ud835\udc5f (4) So, the conclusion here is that equivalent forces are opposite to the reactions because the do the same work. In figure 5 are show some example of equivalent forces where force is concentrated (case a), distributed load (case b) and difference in temperature (case c). Figure 5a. Equivalent force in case where concentrate force acting on bar. Figure 5b. Equivalent force in case where distributed load acting on bar. Figure 5c. Equivalent force in case where difference in temperature is between two sides of bar"
        ]
    },
    {
        "image_filename": "designv11_63_0002900_09544054211023625-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002900_09544054211023625-Figure6-1.png",
        "caption": "Figure 6. Rotary angle measurement with Doppler laser interferometer MCV2002 + RT100.",
        "texts": [
            " Moreover, SGEs of machine spindles in the actual machining will also be taken into account in this study as a main machining error items. It is reported that20 the SGE can be divided into two categories, linear error and rotational error, which includes 21 kinds and 12 kinds of error items, respectively. Here, the LDDM laser measurement system of Doppler laser interferometer MCV2002 according to ISO 230-2(1997) standard52 is adopted to make measurement of the three linear shafts. Doppler laser interferometer MCV2002 with an RT-100 rotator in Figure 6 is used to make measurement of two rational spindles. Then, the software error compensation method with the NC data53 is further exploited for the SGEs compensation. It is worth noting that there are many SGEs from the five-axis hypoid generator and they have obvious effects on the tooth flank topography and performance evaluations,54,55 especially for the actual tooth flank manufacturing geometric accuracy. The detailed measurement and compensation method can refer to the basic methods in Ding et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002295_itaic49862.2020.9339042-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002295_itaic49862.2020.9339042-Figure1-1.png",
        "caption": "Fig. 1. 6-DOF serial robot manipulator with offset wrist [19]",
        "texts": [
            " The upstate function of particle position is same as (7). IV. SIMULATION RESULTS This section presents the simulation result of the 6 -DOF robot manipulator with offset wrist joints performed in MATLAB environment. The standard-PSO and proposed PSO are tested in four cases to prove the increase of the performance of proposed PSO in solving inverse kinematics problem of robot manipulator. The 6 -DOF manipulator with offset wrist joints[19] is selected for inverse kinematics solution, as shown in Fig. 1. Its D-H parameters are shown in Tab.1. The parameters of standard PSO[20] set in this paper is that w = [0.4 ,0.9] and decrease linearly, cl=c2 = 2 .The parameters of Improved PSO is c3 = 0.5,% = 0.8,P=0.01 and others is same as in standard PSO. The stop conditional is maximum iteration which is set as 1 0 0 0 . We randomly selected 4 desired homogeneous matrixes of 6 -DOF end-effector to solve their inverse kinematics problems, they are 0 6 0.27 0.70 0.65 -17.09 0.95 -0.23 -0.15 -47.17 0.04 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002443_012021-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002443_012021-Figure2-1.png",
        "caption": "Figure 2. Support structures and critical overhang angle [6]",
        "texts": [
            " However, there are also design restrictions for the parts produced by LPBF. The rapid heat input and cooling cause the shrinkage of the bonded consecutive layers leading to residual stresses in the part. For overhang features, these residual stresses result in deformation. When the overhang angle is below a critical angle \u03b1crit this deformation can lead to failed built parts caused by high deformation and crash between recoater and part. To avoid it, support structures are used to counteract the stresses and to dissipate heat (figure 2) [6]. To prevent deformation 19th Drive Train Technology Conference (ATK 2021) IOP Conf. Series: Materials Science and Engineering 1097 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1097/1/012021 of the parts after the removal of the support structures, the components might be heat treated after the process, depending on the material. After the LPBF process, the post-processing of the parts on the substrate plate is required. On the one hand, the support structures must be removed mechanically, on the other hand functional surfaces with high surface requirements have to be post-processed by milling or drilling",
            " Series: Materials Science and Engineering 1097 (2021) 012021 IOP Publishing doi:10.1088/1757-899X/1097/1/012021 The most promising approach to fully exploit the technological and economical potential of LPBF is to apply different design strategies like lightweight design, functional integration or improvement, monolithic design, or customized mechanical properties, which are feasible due to the great design freedom. While exploiting the advantages of the technology, the design restrictions that apply to LPBF parts must also be considered. As shown in figure 2, a main consideration when targeting complex parts as a cylinder head are cavities that can be found all over the part. In the first step, the main orientation of the cylinder head towards the build plate is determined, to find an optimal orientation where the number of critical surfaces in terms of LPBF is reduced. The remaining critically surfaces with an overhang angle below the critical angle, need to be addressed. Due to the complexity of the part, many of these overhang-violating surfaces can be found inside the water jackets of the cylinder head as well as in the combustion in-take and out-take"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002263_s00170-020-06477-2-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002263_s00170-020-06477-2-Figure12-1.png",
        "caption": "Fig. 12 Calculation model of multi-tooth cutting force",
        "texts": [
            " and (33) can be obtained F3x \u00bc Fs F3z \u00bc f \u03c6r; \u03b3n; \u03b7s; \u03b7c;\u03b2\u00f0 \u00deF3x \u00f034\u00de F3z is the tangential cutting force, which can be obtained by expanding and simplifying the above equation F3z \u00bc cos\u03b7ssin \u03c6r\u2212\u03b3n\u00f0 \u00de \u00fe tan\u03b2cos\u03b7scos \u03c6r\u2212\u03b3n\u00f0 \u00de cos \u03c6r\u2212\u03b3n\u00f0 \u00de\u2212tan\u03b2cos\u03b7c \u03c6r\u2212\u03b3n\u00f0 \u00de F3x \u00f035\u00de In the cutting force model L03 \u00bc L01L12L23 \u00bc cosicos\u03d5rcos\u03b7s \u00fe sinisin\u03b7s cosicos\u03d5rsin\u03b7s \u00fe sinicos\u03b7s \u2212cosisin\u03d5r sinicos\u03d5rcos\u03b7s\u2212cosisin\u03b7s sinicos\u03d5rsin\u03b7s \u00fe cosicos\u03b7s \u2212sinisin\u03d5r sin\u03d5rcos\u03b7s sin\u03d5rsin\u03b7s cos\u03d5r 2 4 3 5 \u00f036\u00de The force on the shear plane of the cutting region is expressed in the coordinate system S0(X0, Y0, Z0), and the following relation can be obtained F0 \u00bc L03F3 \u00f037\u00de According to the above equation, tangential force Ft, radial force Fr, and axial force Fa are respectively Ft \u00bc F3x cosicos\u03c6rcos\u03b7s \u00fe sinisin\u03b7s\u00f0 \u00de\u2212F3zcosisin\u03c6r Fr \u00bc F3x sinicos\u03c6rcos\u03b7s\u2212cosisin\u03b7s\u00f0 \u00de\u2212F3zsinisin\u03c6r Fa \u00bc F3xsin\u03c6rcos\u03b7s\u2212F3zcos\u03c6r 8< : \u00f038\u00de Generally speaking, the cutting tool used to machine the pinion is only equipped with inner cutting blades and outer cutting blades, and there is no middle cutting blades for slotting. In the process of tooth generating, the cutting tool has the situation that single cutting blade and multiple cutting blades work alternately, so it is necessary to analyze the force in each stage of the tooth cutting process. The cutting force calculation model of multi-cutting blade machining is established, as shown in Fig. 12. If the feeding speed of the cradle is set as Vp (mm/min), then, the angle and time required for any cutting blade tomove from the current position to the position of the adjacent cutting teeth are \u03c6c \u00bc 2\u03c0 Zp ; tc \u00bc r1\u03c6c Vp \u00f039\u00de According to the relevant knowledge of solving the intersection line of the cutting tool tip and the outer circle of the wheel blank end face, the cradle rotation angle qs and the cutting tool rotation angle \u03c6cs at the beginning of their contact can be calculated. The resultant of the tangential cutting force of each cutter tooth is Ftx \u00bc \u2211 i\u22120 N1 Fto i\u00f0 \u00decos\u03d5c i\u00f0 \u00de \u00fe \u2211 j\u22120 N2 Fti i\u00f0 \u00decos\u03c6c j\u00f0 \u00de Fty \u00bc \u2211 i\u22120 N1 Fto i\u00f0 \u00desin\u03c6c i\u00f0 \u00de \u00fe \u2211 j\u22120 N2 Fti i\u00f0 \u00desin\u03c6c j\u00f0 \u00de Fres \u00bc Ftx 2 \u00fe Fty 2 1=2 \u00f040\u00de where \u03c6cs \u2264\u03c6c \u2264\u03c6ce, i, and j cannot be equal to zero at the same time, because at least one of the cutting teeth is in cutting state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002307_012011-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002307_012011-Figure2-1.png",
        "caption": "Figure 2. Solid models of the analysed gears.",
        "texts": [
            " The gears have a constant torque loads T1 and T2. Stiffness of meshing k(t) is time-varying and constitutes the main reason for gear vibrations. Two tooth stiffness models were analysed. The first one according to the ISO 6336 -1 standard [21], whereas the second one progresses similarly to the real process according to Cai [7]. Damping c in the meshing was assumed to be constant. Meshing friction was analysed for four variants of the friction model. Solid models of the gears subjected to analysis were made (Figure 2). On their basis the moments of inertia were determined, assuming steel as the material. The remaining parameters were shown in Table 1. CMES 2020 Journal of Physics: Conference Series 1736 (2021) 012011 IOP Publishing doi:10.1088/1742-6596/1736/1/012011 The equations of motion were introduced on the basis of the free body diagram as well as the Newton\u2019s second law. The forces resulting from stiffness and damping act in the direction of a line of action (LOA). This line is tangential to the base circles rb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001399_s00170-020-05913-7-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001399_s00170-020-05913-7-Figure17-1.png",
        "caption": "Fig. 17 Forming process of asymmetric spinning using the rotational pass set",
        "texts": [
            " Pass angle \u03b8 is the angle of the endpoint relative to the start point of a pass. It decreases by an incremental angle \u0394\u03b8 for each pass relative to the last pass. The linear pass is used at the end of the process to push the wall on the mandrel. The position of the roller is controlled synchronously in both radial and axial directions during every revolution of the spindle to achieve the asymmetric shape as illustrated in Fig. 16. The forming process using the rotational pass set in asymmetric spinning is shown in Fig. 17. The inclination angle \u03c6 is determined by the difference of half cone angle on each side of the product (\u03c6 =\u03b1L \u2212\u03b1U). The product is formed from a sheet blank to shape (a) with an inclination angle \u03c6. Then, the shape of the product is changed to (b) by keeping \u03c6 constant. Next, \u03c6 is reduced to zero by achieving shape (c). Finally, a linear pass is used to force the wall of the product onto the mandrel. The experiments on cylindrical cups using the rotational pass set in asymmetric spinning were conducted with the parameters given in Table 2. For comparison, cylindrical cups using the same pass set in conventional spinning were formed. An example of the formed product using the rotational pass set is shown in Fig. 18. The thickness on the upper and lower sides of the wall (tU and tL shown in Fig. 17) were measured axially. Figures 19 and 20 show the axial distribution of wall thickness ratios (tU/t0 and tL/t0) with different parameters of \u0394\u03b8 = 3\u00b0, 4\u00b0, 5\u00b0, 6\u00b0 mm and \u03c6 = 6\u00b0, 24\u00b0, 48\u00b0, 72\u00b0. The dotted lines are distributions of thickness ratio (tC/t0) in conventional spinning (when \u03c6 = 0\u00b0). The distributions of tU/t0 and tL/t0 are similar to tC/t0. The difference between tU/t0 and tL/t0 is caused by the varying half cone angle of the product during forming. The difference is much smaller compared with that in the translational pass set and it increased slightly as \u03c6 increased as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001249_s00170-020-05744-6-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001249_s00170-020-05744-6-Figure7-1.png",
        "caption": "Fig. 7 Isometric view of smallscale and large-scale actuators",
        "texts": [
            " The chitosan films are attached to the actuator body using 3D printed transversal members that span the width of the actuator and can accommodate fasteners to clamp the film. This complex-geometry assembly was designed following Design for Assembly/ Manufacturing principles (DFA/DFM). Accordingly, the parts were dimensioned with specific tolerances following limits and fits theory to avoid room for error in mating parts while still accommodating room for assembly. Both small-scale (~ 250 mm long) and large-scale (~ 2 m long) cantilever actuators were designed using the above said design methodology as shown in Fig. 7. The small-scale actuators were used to test the overall functioning and assembly logic of the system, while the large-scale actuator served as a diagnostic test of the limits of the assembly with higher forces involved. To predict the shape change of the large-scale actuator, the design was fed into the Kangaroo physics engine in Rhinocerous. The shape change simulation was accomplished by instilling geometric changes in the model based on the characterized expansion of the chitosan films. As the largescale prototype was a spanning cantilever, it was evident that the joints toward the fixed end experienced the most load during the motion of the structure, and if the design were to scale up further, these forces would significantly increase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002250_s12666-020-02152-y-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002250_s12666-020-02152-y-Figure3-1.png",
        "caption": "Fig. 3 Casting, mold and methoding (gating and feeding) elements",
        "texts": [
            " There is another exacerbating factor. Casting lifecycle has four major stages, which are usually carried out in separate organizations: component design at original equipment manufacturer (OEM); pattern development in tool-room; process planning, casting and fettling in foundry; and component machining in another company. Process planning (also called methoding) includes the design of gating system to lead molten metal into mold cavity and feeding system to compensate solidification shrinkage (Fig. 3). The component design is modified during tooling and methoding to improve its manufacturability. This can increase its weight (typically 5\u201315%), which has to be either accepted by the OEM or machined off, implying higher cost. The above factors along with metal and process also affect overall yield\u2014the ratio of component weight to poured weight, which can range from 40 to 80%. While remelting and recycling are an integral part of metal casting, it also implies preventable loss of energy and productivity",
            " They also need to modify the methoding design in a CAD program, transfer the casting model (along with mold, core, feeding and gating elements) to a simulation program, compute the results, visualize the defects, check if they are still present inside the casting and repeat the steps till the desired quality and optimal yield are achieved. This is cumbersome for those who have limited knowledge, skills and patience for complex software programs. The obvious solution is to integrate casting design and simulation in a single system, and incorporate intelligent algorithms to quickly \u2018get the job done\u2019 with minimal user inputs. Methoding of castings is, however, quite complex. It includes many inter-connected decisions regarding mold cavities, gating and feeding (Fig. 3). The number and layout of cavities are optimized to achieve maximum productivity (ratio of cast metal to mold material per cycle) while minimizing mold damage and casting defects [12]. The gating system comprises pouring basin, sprue, well, runner, extension, gates and filters. Their number, location, shape and dimensions, along with pouring parameters (temperature, pressure, time and rate) are designed to ensure complete, uniform and smooth filling of mold cavity with clean metal. For this purpose, good insights can be derived from experimental studies with water in transparent molds [13]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure52.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure52.1-1.png",
        "caption": "Fig. 52.1 Concepts Ideation",
        "texts": [
            " The spokes length can be adjusted by attaching or removing parts of the spoke. Spokes of one side of the wheel can be shortened and the other side lengthened to negotiate the steepness of the slope so as to maintain the machine in a horizontal position. The scraper is attached on the rear side of the machine between the main body and the handles of the machines. This scraper will be the unit in contact with the soil for cutting the bench terrace. The machine is intended to be hand-guided by an operator as shown in Fig. 52.1a. In this approach, the idea is to pave a pathway for the machine to manoeuvre over hilly terrains by cutting the slopes and forming terrace fields. The system consists of five sub-units, viz., the engine, the chassis or the body, a power take-off (PTO) unit, a rotary cutting unit and a conveyor belt unit (Fig. 52.1b). The rotary blade unit is placed in front of the engine connected by a power take-off unit for transmitting power from the engine to the rotary cutting unit. The vertically height adjustable rotary cutting unit will be positioned close to the surface of the hill/mountain slope in order to cut the soil. The falling debris or the loose soil which are cut out will fall on the conveyor belt unit. They will then be transferred for filling up to the other end of the machine which is at the lower level area of the slope. This approach will provide a stable and safe area for the machine and the operator to work along the hilly terrains. This concept consists of four sub-units, viz., the engine, the chassis or the body, a power take-off (PTO) unit, a chain drive cutting unit as shown in Fig. 52.1c. In this concept, power from the engine of the power tiller is transferred to the chain drive system by the PTO unit. The cutting blades will be mounted along the length of the chains which in turn are mounted on the sprockets and will cut the soil when the chain rotates. The middle pair of sprockets acts as the pivot about which the cutting unit of the chain drive system rotates so that the height of cutting operation can be adjusted. This concept proposes a rotary cutting unit with its own separate engine attached to the front of a power tiller as shown in Fig. 52.1d. This concept is designed to provide movement of the rotary cutting unit along the vertical and horizontal directions and also provide the ability to cut the soil at a certain angle away from the normal position of the rotary unit. The operator/handler will walk behind the power tiller, so the thrown off soil will not affect the operator. The cutting blades were designed to \u201cthrow off\u201d to the cut soil opposite to the direction of the cutting action. The selection of the most appropriate concept is based on three levels; discussing the advantages and disadvantages of each concept, receiving inputs and opinions from individuals familiar with design processes, decision matrix to rate each concept"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002120_s00170-020-06571-5-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002120_s00170-020-06571-5-Figure6-1.png",
        "caption": "Fig. 6 Coordinate system of wheel and pinion engagement",
        "texts": [
            " rWline \u00bc xW hc; \u03b8c\u00f0 \u00de yW hc; \u03b8c\u00f0 \u00de zW hc; \u03b8c\u00f0 \u00de 1 T \u00f08\u00de Because the unit tangent vector tWline along the wheel contact path should be the direction along which the instantaneous contact point moves on the wheel flank, it can be expressed as follows: tWline \u00bc drWline dhc = drWline dhc \u00f09\u00de where drWline dhc \u00bc \u2202xW \u2202hc \u00fe \u2202xW \u2202\u03b8c \u22c5 d\u03b8c dhc iW \u00fe \u2202yW \u2202hc \u00fe \u2202yW \u2202\u03b8c \u22c5 d\u03b8c dhc jW \u00fe \u2202zW \u2202hc \u00fe \u2202zW \u2202\u03b8c \u22c5 d\u03b8c dhc kW \u00f010\u00de where iW, jW, and kW are the unit vectors along the xW-, yW-, and zW-axis of the wheel-moving coordinate system oW \u2212 xWyWzW (shown in Fig. 6), respectively. According to the above method, for the wheel convex flank, the first-order parameters of each point on the contact path, and the first-order and second-order parameters of the mean contact point are obtained, as well as the unit tangent vector along the contact path at the mean contact point. For the wheel concave flank, the first-order and second-order parameters of the mean contact point are determined as well. In Fig. 6, oM \u2212 xMyMzM is the fixed coordinate system of the wheel and pinion engagement, and oW \u2212 xWyWzW is the wheelmoving coordinate system connected rigidly with the wheel. The pinion-moving coordinate system is represented by oP \u2212 xPyPzP connected rigidly with the pinion, and oPd \u2212 xPdyPdzPd is the pinion-fixed coordinate systemwithout rotation with the pinion. The xW-axis and xM-axis coincide with the wheel axis, xPaxis and xPd-axis coincide with the pinion axis, point oW and point oM coincide with the wheel crossing point, and point oP and point oPd coincide with the pinion crossing point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure19-1.png",
        "caption": "Fig. 19. Proposed HEBFM prototype of (a) rotor, (b) stator, (c) top side, and (d) bottom side.",
        "texts": [
            " Downloaded on June 05,2021 at 19:52:57 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (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. A prototype of the proposed HEBFM is manufactured using the parameters in Table I to evaluate the proposed design based on the winding factor modulation method. Main components of the tested HEBFM are shown in Fig. 19. The photograph of the test bench is shown in Fig. 20. The operating conditions in experiment are set to be the same as those in simulation. The experiments are conducted with excitation currents of 5A, 0A, and -5A (i.e., IDC=5A, IDC=0A, and IDC=-5A), respectively. For sake of fair comparisons, the analytical, simulation and experiment result of Back EMF in the three cases are provided in Table X. FEA and measured Back EMFfield current curve are shown in Fig. 21. With the field current changes from -7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure26-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure26-1.png",
        "caption": "Fig. 26 Thickness analysis of the upper panel",
        "texts": [
            " After the forming process in the first study on the upper panel, a maximum of 37% thinning was observed. Subsequently, the sharp corners were rounded in areas where there was substantial thinning, leading to a decrease in the maximum level of thinning to 22%, as shown in the figure below. Another issue that prevents production of the part is wrinkles in the middle of the part. The crimping process was implemented in the design to address this problem. The thickness analysis of the upper panel is shown in Fig.\u00a026. After the forming process in the first study on the lower panel, a maximum of 15.6% thinning was observed. To take precautions against the shear around sharp corners and to increase the stiffness of the part, the crimping process was implemented. Additionally, some radiuses were increased from 3 to 5\u00a0mm, which had an impact on thinning. The thinning analysis of the lower panel is shown in Fig.\u00a027. Main aim of this study was to reduce weight of engine cross members by considering stiffness and stress levels in light commercial vehicles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001719_biorob49111.2020.9224276-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001719_biorob49111.2020.9224276-Figure3-1.png",
        "caption": "Fig. 3: Application of the handler in-plane displacement to manually move the ROI where is estimated the elastogram in the ultrasound image.",
        "texts": [
            " On the other hand, the haptic control mode (presented in section III-B) generates a force from the estimated elastogram of the observed tissue and render it through the haptic device to the user. This elastogram is estimated in a chosen region of interest (ROI) of the ultrasound image. We will not recall here the principle of the elastogram estimation since it is not the topic of this paper but we will consider in the next that it is provided by an image processing as the one we detailed in [11]. The teleoperation mode consists in applying to the probe frame Fcp the relative displacement of the haptic handler frame Fh (see Fig. 3) that results from the user manual motion. To measure this relative displacement, we first define the initial and the current poses of the handler with respect to the base frame Fb of the haptic device as b Ph0 \u2208 SE(3) and b Ph \u2208 SE(3), respectively. The initial pose of the handler is measured only at the initialization time when the user launched the teleoperation mode and its current pose is measured at every time. In the next, we will use the notation a Mb to refer to the 4\u00d74 homogeneous matrice that describes a relative pose a Pb of a frame Fb with respect to a frame Fa",
            " (14) with \u03b5 being the scalar strain value estimated from the ROI, and it gives: k = AE L (17) where A is the region area of the virtual probe that senses the local strain of the tissue in the ROI and that corresponds in our case to the elliptical surface A = \u03c0\u03c3x\u03c3y of the Gaussian mask Gm. E is the Young\u2019s modulus that we set to the value E = 3kPa of healthy tissue and L is the original length of the virtual spring that corresponds to the width d of the rectangular contact surface between the real ultrasound probe and the tissue (see Fig. 1b). 2) Virtual probe control and force feedback: Fig. 3 illustrates the principle that consists in moving the ROI to follow the displacement of the user measured by the handler of the haptic device. If the user applies motion at the handler of the haptic device Fh, then the center of the ROI (uc, vc) is shifted with a displacement \u2206d \u2208 R 2 proportional to the displacement of the handler such that, \u2206d = S\u2206h, (18) where \u2206h \u2208 R 2 is the in-plane relative motion applied to the handler that is directly measured from the x and y translation components of the 4th column of the homogeneous matrix h0 Mh introduced in eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000627_uemcon47517.2019.8993050-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000627_uemcon47517.2019.8993050-Figure5-1.png",
        "caption": "FIGURE 5. Sectorized 3D antenna gain pattern assumed for the transmitters and receivers in the wearable network.",
        "texts": [
            " For example, this could be the maximum transmit power of devices as set out by regulations, in which case this would be a worst case scenario. The noise power normalized by the signal power observed at a reference distance is denoted as \u03c3 2. The transmitters and receivers are assumed to have directional antennas. The antenna array pattern is characterized by four parameters - the azimuth beam-width \u03b8a, the beam-width of the antenna main-lobe in the elevation \u03b8e, the main-lobe gain G within the beam-width, and the side-lobe gain g outside the main-lobe as shown in Fig. 5. We use subscripts t and r to denote the antenna parameters at the transmitter and the receiver, respectively. The transmitter main-lobe of Ti is assumed to be pointed in the direction given by the azimuth angle \u03c6a i and elevation \u03c6 e i . In this model, omni-directional transmission and reception is a special case with G = g = 1. The receiver-transmitter pair associated with user i are assumed to be aligned. From the reference receiver\u2019s perspective, this assumption leads to a random transmit gain for the interference from Ti as seen at R0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000284_j.compeleceng.2019.106534-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000284_j.compeleceng.2019.106534-Figure10-1.png",
        "caption": "Fig. 10. Magnetic field of motor under load. (a) Bs0 = = 0 mm (b) Bs0 = = 0.4 mm (c) Bs0 = = 0.8 mm (d) Bs0 = = 1.8 mm.",
        "texts": [
            " This means that reducing the slot width of the rotor can effectively decrease the torque ripple, because the magnetic permeability of the closed groove is larger, the torque ripple can be more effectively eliminated. Similarly, when the load torque is increased to 50 and 60 N m, the motor torque ripple of the closed slot is also reduced respectively by 3.52% and 3.78%, respectively, compared with those with 1.8 mm as the open slot. It demonstrates that appropriate reduction of the width of the rotor slot can effectively suppress the motor torque fluctuation. Fig. 10 shows the magnetic flux density distribution of the BIM under load state. The magnetic field distribution under different slot widths is similar. The magnetic field density in the stator yoke is relatively dense and in the rotor yoke is sparse. However, in rotor tooth, there exists partial magnetic saturation. In Fig. 10 (a), as the rotor slot width is 0 mm (closed slot), the magnetic field is saturated in many parts of the rotor tooth, and the maximal flux density is 3.17 T. When the rotor slot width is 1.8 mm, the saturation only occurs in two places of rotor tooth, and one is 2.54 T, while another is 1.97 T. This indicates that the flux saturation in the rotor tooth is eased with the increase of the slot width. However, this is not applicable to the stator tooth. By comparing the magnetic density of stator tooth with different slot widths, it can be found that excessive slot width will aggravate the stator flux saturation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure16-1.png",
        "caption": "Fig. 16 The robot manipulator axes.",
        "texts": [
            " By means of the pins placed in these limitations, the ROM of the joints can be limited to the desired values. The kinematic and dynamic analysis of a robotic system is important for robot control. When performing kinematic and dynamic analysis of robots, classical manual calculation techniques can be used. However, as the DOF of the robot increases, these calculations become more complicated. For this reason, for three-DOF PHYSIOTHERABOT/w1, the analysis programs that can calculate the parameters related to the system model were used. Fig. 16 shows the axes of the robot manipulator for kinematic analysis. The Denavit-Hartenberg parameters according to the axes of the robot manipulator shown in Fig. 16 are given in Table 6. 431Impedance control applications in therapeutic exercise robots transformation matrices are as follows: T0 1 \u00bc cos\u00f0q1\u00de 0 sin\u00f0q1\u00de 0 sin\u00f0q1\u00de 0 cos\u00f0q1\u00de 0 0 1 0 0 0 0 0 1 2 664 3 775 (96) T1 2 \u00bc sin\u00f0q2\u00de 0 cos\u00f0q2\u00de 0 cos\u00f0q2\u00de 0 sin\u00f0q2\u00de 0 0 1 0 0 0 0 0 1 2 664 3 775 (97) T 2 3 \u00bc cos\u00f0q1\u00de sin\u00f0q3\u00de 0 l1 cos\u00f0q3\u00de sin\u00f0q1\u00de cos\u00f0q3\u00de 0 l1 sin\u00f0q3\u00de 0 0 1 l2 0 0 0 1 2 664 3 775 (98) 432 Erhan Akdogan and Mehmet Emin Aktan The drawings of the manipulator were made by the solid modeling program. The link parameters obtained from these drawings are given in Table 7",
            " The impedance parameters Md, Bd, Kd, and S are diagonal matrices for which the first three entries are related to the translational movement of the end effector on x0, y0, and z0 axes and the last three entries are related to the rotational movement of the end effector around x0, y0, and z0, respectively. For example, in order to prevent the end effector move on x0 axis, the first entries of the impedance parametersMd, Bd, andKd are selected asmh, bh, and kh. The first entry of the switching matrix S as position-based impedance control (s \u00bc 1), and the desired position for this axis is 0 degree. Flexion-extension: During the flexion-extension movement, the origin of the coordinate system of the end effector moves in the y0 and z0 axes and rotates about the x0 axis (see Fig. 16). As a result, impedance parameters with corresponding axes are selected as the desired impedance parameters whereas the others are selected as the hold impedance parameters to avoid possible radial-ulnar deviation and pronation-supination movements. Md \u00bc diag\u00bd100, m, m, m , m , 1 Bd \u00bc diag\u00bd1000, b, b, b , b , 10 Kd \u00bc diag\u00bd10,000, k, k, k , k , 100 S \u00bc diag\u00bd1, s, s, s, 1, 1 435Impedance control applications in therapeutic exercise robots The operator diag[x1, x2, \u2026, xn] denotes a block diagonal matrix whose elements on the main block diagonal are x1, x2, \u2026, xn"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001435_10.0001388-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001435_10.0001388-Figure2-1.png",
        "caption": "Fig. 2. Airflow past an ascending football with a positive pitch. The air drag can be characterized by a single force applied at the \u201ccenter-of-force\u201d point that leads to an aerodynamic torque s about the ball\u2019s center of mass. If v and the symmetry axis s\u0302 lie in the plane of the page, then s points into the page. Air streamlines are indicated by the curved lines. The ball will wobble if L, v, and s\u0302 are not collinear.",
        "texts": [
            " The simplest attempts to explain the tangency of the ball\u2019s axis and its velocity, however, are not satisfactory. The dynamics of rapidly spinning tops don\u2019t provide obvious analogs, nor does a weathervane analogy, with the on-rushing air serving to point it into the \u201cwind.\u201d The simplest version of this explanation fails since the football lacks the front-back asymmetry of a weathervane. Upon careful consideration of the effect of the air acting on the ball, a more serious question arises, as illustrated in Fig. 2. The attitude and flight of the ball can be fully characterized by three vectors: its CM velocity v, the unit vector aligned with its long axis s\u0302, and its angular momentum L; for definiteness, we choose s\u0302 to be in the direction such that the angle between it and L is acute. For a \u201cperfect\u201d spiral pass, v, s\u0302, and L are initially co-linear. The ball will wobble if (a), L and s\u0302 are misaligned, causing s\u0302 to execute torque-free precession about L, or (b), L and s\u0302 are closely aligned, but not with v, causing a nutation akin to that of a rapidly spinning top",
            " The forces that the air exerts on the moving ball are complicated; Reynold\u2019s numbers associated with the ball\u2019s CM motion and its spin indicate that the flow of air over the ball\u2019s surface is largely turbulent. Based on the wind tunnel tests of Refs. 5 and 6, however, these forces can be described heuristically and resolved into a single net force acting through the ball\u2019s CM and a force couple that exerts a pure torque about the CM. This couple can further be resolved into a \u201cspindrag\u201d torque along s\u0302 and a second torque s that acts about the CM in a direction perpendicular to s\u0302. The latter torque is the result of the \u201cangle of attack\u201d between s\u0302 and v\u0302; as indicated in Fig. 2. In our analysis, we neglect the spin-drag torque. We also ignore the net drag force on the ball, so that we take the ball\u2019s CM position to be determined solely by its kinematic initial conditions and the force of gravity. Thus, v will always be tangent to a parabolic trajectory lying in the xz plane. The only effect of air drag in our calculations is, thus, to change the ball\u2019s attitude, i.e., its pitch and yaw; these angles are all we need to understand the trajectory-following phenomenon and to resolve the paradox"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001142_ilt-12-2019-0542-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001142_ilt-12-2019-0542-Figure3-1.png",
        "caption": "Figure 3 Rigid single mass bearing-rotor system",
        "texts": [
            " Obtaining partial derivatives of small perturbations D\u00ab , \u00ab0Du , D \u00ab , \u00ab0D u , Dz and D z from equation (1), and six coupled dynamic Reynolds equations including thermal effect and small disturbance. After receiving disturbance pressure in six directions by FEM, the stiffness and damping coefficients are obtained via Simpson integral. Basis upon static equilibrium calculation, the bearing-motor system can be simplified as a rigid single mass bearing-rotor system, and the oil film acts as a combination of spring and damping shown in Figure 3. Then the system dynamics equation is established which takes the thermal effect into account, the dimensionless form (Guo et al., 2019) can be expressed as: k 3 eq 1A k2 eq 1B keq 1C keqg 2 st 1D g2 st 1E \u00bc 0 F k2 eq 1 I keq 1 J g2 st 1K \u00bc 0 8>< >: (9) where the coefficients A K are linear combination of stiffness and damping coefficients. So the dimensionless critical mass is as follows: Mst \u00bc keq g2 st (10) Obtain keq and g2 st by solving equation (9) with Newton iterative method, then acquire the dimensionless critical mass in accordance with equation (10)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001521_1754337120953005-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001521_1754337120953005-Figure2-1.png",
        "caption": "Figure 2. Definition of the spin axis. The spin axis was defined as the angle between the Z-axis and the projected angular velocity vector.",
        "texts": [
            " A 2D coordinate of the marker attached to the seam of the ball and four 2D coordinates on the outer circumference of the ball, which were recorded using the camera behind the player (Camera 1 in Figure 1), were read during a recognizable frame (almost one half-revolution). Then, the angular velocity of the ball was calculated using the computer vision method developed by Murata.20 TrackMan showed the spin axis as twodimensional. Therefore, the angular velocity vector was converted to the projected angle in the XZ-plane, and the spin angle was defined as the angle between the Zaxis and the angular velocity vector (Figure 2). Pearson\u2019s product-moment correlation coefficient was calculated to check the agreement between the speed, spin rate, and direction measured using each method. Furthermore, the inter-method reliability was evaluated using Bland\u2013Altman analysis21 and intraclass correlation coefficients (ICCs). The ICCs were evaluated in accordance with a previous study.22 Statistical analysis without outliers was performed because the primary purpose of this study was to verify the accuracy of TrackMan for research purposes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000573_012007-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000573_012007-Figure6-1.png",
        "caption": "Figure 6. The relationship between transverse vibration velocity of the stator and steady-state speed of the slider in one period",
        "texts": [
            " The threshold of the relative displacement for the two phases is denoted by mu and the interface friction forces are as follows , ( ) , ( ) , , , ( ) , ( ) k N b s r m f k r k N b s r m F w t v u t u F c u F w t v u t u            (19) ( ) , ( ) , ( ) , ( ) , ( ) , ( ) b s p N b s r m f b s p N b s r m w v F w t v u t u F w v F w t v u t u                   (20) in which  is the viscous friction coefficient and sv is the steady-state speed of the slider. And k is the friction coefficient in the stick phase that is proportional to the magnitude of relative displacement. And p is the dynamic friction coefficient. The proportional coefficient c is given by /p mc u . Figure 6 shows the relationship between ( )bw t and sv in one period. ISPECE 2019 Journal of Physics: Conference Series 1449 (2020) 012007 IOP Publishing doi:10.1088/1742-6596/1449/1/012007 The motion equation of the slider is obtained by force equilibrium as follows ,s r s d g load dv m c v F F F dt     (21) sign( ) ( 9.81 ),g s sd pF v F m   (22) where , ,r sdm c  are the mass of slider, viscous damping coefficient, and dynamic friction coefficient of linear guide and ,d loadF F are the driving the load forces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000882_j.cagd.2020.101870-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000882_j.cagd.2020.101870-Figure1-1.png",
        "caption": "Fig. 1. Notation for sides and angles.",
        "texts": [
            " E-mail address: izmestiev@dmg.tuwien.ac.at. https://doi.org/10.1016/j.cagd.2020.101870 0167-8396/\u00a9 2020 Elsevier B.V. All rights reserved. congruence, there is a one-parameter family of quadrilaterals; in other words, a four-bar linkage has one degree of freedom. It follows that there is a relation between any pair of angles in the quadrilateral. Also, it should be possible to express all four angles as functions of a single variable. Denote the side lengths and the angle values of a quadrilateral as shown in Fig. 1. The relation between two opposite angles can be found by expressing in two ways the length of the diagonal which separates these angles, see Fig. 2. Namely, by the law of cosines one has a2 + d2 + 2ad cos\u03d5 = e2 = b2 + c2 + 2bc cos\u03c8, which implies 2ad cos\u03d5 \u2212 2bc cos\u03c8 + (a2 \u2212 b2 \u2212 c2 + d2) = 0. (1) This linear equation between the cosines of two opposite angles implies that there is a linear dependence between the cosines of the input and the output angles in every scissors linkage, see Fig. 3. In particular, if the input and the output angles coincide in two linkage configurations, then they coincide in all configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002753_s00202-021-01301-w-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002753_s00202-021-01301-w-Figure1-1.png",
        "caption": "Fig. 1 Segmented magnets in rotor of PM BLDC motor",
        "texts": [
            " Eddy current in proportion to the length size of the motor flows in these magnets of radial structure. The most important effect of eddy current is heating. This heating has an effect on the 1 3 performance parameters of the motor. One way to decrease eddy current losses is reducing the current in which it runs. Segmented structure of permanent magnets in axial direction is for evaluating the effect of eddy current losses. Segmented structure in axial direction of magnets that make up the rotor of in-wheel BLDC motor is shown in Fig.\u00a01. The magnets are divided into blocks which are isolated from one another electrically. They can divide in the axial direction with at least two blocks per pole. Effective resistance is increased when eddy-current paths are divided by segmentation into smaller loops. Because of the eddy-currents are resistance-limited, the associated power loss is decreases [9, 11]. Decrease in losses as a result of segmented structure of PM\u2019s in axial direction can be expressed by Eq.\u00a0(1) [9, 30, 31]. variable a and b are the half of the width and the half of the length of the conductor, \u03b4 is the skin depth; d is the thickness; m is the number of series solutions; H is the imposed harmonic magnetic field; and \u03bbm, \u03b3, \u03b2m, \u03b2mr, and \u03b2mi are the coefficients determined by \u03b4 and m [12]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003191_itec51675.2021.9490116-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003191_itec51675.2021.9490116-Figure2-1.png",
        "caption": "Fig. 2. Drive system configuration of the proposed five-phase synchronous reluctance motor.",
        "texts": [],
        "surrounding_texts": [
            "reluctance motor is able to offer a non-permanent magnet alternative. This paper proposes the a robust non-permanent magnet five-phase synchronous reluctance motor. The torque capability of this type of motor can be improved by utilizing harmonic current injection. The machine topology, system configuration, rotor structure selection are illustrated and discussed. The electromagnetic performance and effect of the harmonic current component are verified by finite element analysis, which shows that the torque capability of the proposed five-phase synchronous reluctance motor can be significantly improved. Additionally, the proposed reluctance motor exhibits a robust structure and is suitable for high-speed and hightemperature operation.\nKeywords\u2014Synchronous reluctance motor, five-phase,\ntraction motor, permanent magnet-free, electric vehicles.\nI. INTRODUCTION\nInterior permanent magnet (IPM) machines utilizing high strength rare-earth magnets have been dominant in industrial drives, electric vehicles drivetrain, and renewable energy systems [1]-[18]. Various control schemes are also proposed to improve the performance of drive systems [19]-[30]. Unfortunately, the rare-earth magnets face price fluctuation, limited resources, and uneven distribution around the world. As a result, it is important to develop the non-rare-earth and/or non-PM alternatives [31]-[36].\nOne of the straightforward alternatives is a PM-assisted synchronous reluctance motor since the conventional IPM machines are actually a combination of PM machine and synchronous reluctance machine. This type of machine uses low-cost ferrite magnets, instead of rare-earth magnets. The overall performance has been investigated and evaluated by several researchers, which shows that the PM-assisted synchronous reluctance motor can be competitive to the rareearth IPM machines in terms of power/torque capability and efficiency [37]-[47]. However, there is still a potential demagnetization risk for ferrite magnets due to their relatively lower coercivity. Induction machines, on the other hand, are the most popular AC machines in general-purpose industrial drives. Induction machines are capable of producing electromagnetic torque without any types of permanent\nmagnets, which is a low-cost and reliable candidate for traction applications [48]-[50]. Furthermore, the copper rotor was proposed to improve efficiency. However, the efficiency of induction machines is still can not be competitive to IPM machines, particularly under low-speed operation. Synchronous reluctance machines are an appealing candidate to eliminate permanent magnets due to their low-cost rotor structure, negligible rotor loss, synchronous operation feature, and simple control implementation. Synchronous reluctance machines are considered as an alternative to replacing induction machines since the 1980s. However, the major drawbacks of conventional three-phase synchronous reluctance machines include low torque density and poor power factor.\n978-1-7281-7583-6/21/$31.00 \u00a92021 IEEE\n20 21\nIE EE\nT ra\nns po\nrt at\nio n\nEl ec\ntr ifi\nca tio\nn Co\nnf er\nen ce\n& E\nxp o\n(IT EC\n) | 9\n78 -1\n-7 28\n1- 75\n83 -6\n/2 1/\n$3 1.\n00 \u00a9\n20 21\nIE EE\n| D\nO I:\n10 .1\n10 9/\nIT EC\n51 67\n5. 20\n21 .9\n49 01\n16\nAuthorized licensed use limited to: Hoseo Univ. Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply.",
            "The purpose of this paper is to present the design and performance analysis of a five-phase synchronous reluctance motor for torque capability improvement. The effect of the third harmonic current component can be utilized in the fivephase synchronous reluctance motor. The machine topology and drive system will be introduced first. One of the key issues of the proposed motor is the selection of rotor structure, which will be discussed in section two. Finally, the overall electromagnetic performance and effect of the third harmonic current component will be verified by finite element analysis. It will be shown that the torque density can be improved by utilizing the harmonic current injection scheme and the proposed five-phase synchronous reluctance motor is an appealing non-PM candidate for taction applications.\nThe machine topology and drive system configuration of the proposed five-phase synchronous reluctance motor drive is introduced in this section. Fig. 1 shows cross-section, power converter, and drives topologies for the proposed fivephase synchronous reluctance motor. The five-phase symmetric distributed windings and simple salient reluctance rotor are employed in the proposed synchronous reluctance motor. The rotor structure selection will be discussed in the next section, which will show that the simple salient reluctance rotor is more suitable for the proposed five-phase synchronous reluctance motor because of its good coupling capability between fundamental and third harmonic MMF [51]-[52]. In addition, a full-bridge five-phase power electronic converter was employed to achieve vector control and harmonic current injection for proposed five-phase synchronous reluctance motor drives. The third harmonic current component can be utilized for average torque production.\nThe electromagnetic torque of the proposed five-phase\nsynchronous reluctance motor can be expressed [53]\n(1)\nwhere P is pole numbers, Ld1, Lq1 are fundamental d-axis and q-axis inductance, Ld3, Lq3 are third harmonics d-axis and qaxis inductance, id1, iq1 are d-axis and q-axis fundamental current, id3, iq3 are d-axis and q-axis third harmonic current. The first term of this equation represents the electromagnetic torque produced by fundamental current, which is similar to the torque production in a conventional three-phase synchronous reluctance motor. In addition, the third term represents electromagnetic torque produced by the third harmonic current only, which has a very small contribution and can be neglected. It should be noted that the second term indicates the interaction of fundamental and third harmonics current and MMF. This interaction can produce useful\naverage torque because of the special coupling effect of the salient reluctance rotor. The torque capability can be improved and enhanced by utilizing the third harmonic injection.\nIII. DESIGN AND VERIFICATION\nA. Rotor topology selection\nRotor topology or structure plays a key role in order to improve additional torque capability in the proposed fivephase SynRM. Generally, the topologies of reluctance rotor can be classfied into three types, including salient pole, multilayer flux barriers, and axially laminated rotor. Both multilayer flux barriers and axially laminated rotors can achieve a higher saliency ratio than simple salient pole rotors. The multi-layer flux barrier reluctance rotor is an attractive option in conventional three-phase synchronous reluctance motor since it compromises saliency ratio and manufacturing complexity. However, in order to utilize the effect of the third harmonic current and MMF, the simple salient rotor is more suitable in the proposed five-phase SynRM. The influence of different rotor structures on the torque capability of fivephase SynRM has been investigated in [49]. In addition, the simple salient reluctance rotor is also attractive for high-speed traction applications. Therefore, the simple salient reluctance rotor is employed in the proposed design and analysis to meet electromagnetic and mechanical design requirements.\nB. Performance verification\nIn this section, the finite element analysis was employed to predicted electromagnetic performance and characteristics of\n1 1 1 1 13 1 3 1 3\n3 3 3 3\n( ) 2 ( )5\n3( )2 2\nq d d q q d d q\ne\nq d d q\nL L i i L i i i iP T\nL L i i\n\u2212 \u2212 \u2212   =  \n+ \u2212  \nAuthorized licensed use limited to: Hoseo Univ. Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_63_0001053_iceee49618.2020.9102576-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001053_iceee49618.2020.9102576-Figure1-1.png",
        "caption": "Figure 1. Space-vector diagram of IPMSM",
        "texts": [
            " Section 3 explains self-commissioning scheme which is consisting of RLS technique, parameters identification RLS models and automatic tuning of both current and speed controllers\u2019 gains. Experimental results are provided in Section 4 and Section 5 concludes the paper. 338 978-1-7281-6788-6/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: University of Exeter. Downloaded on June 22,2020 at 07:47:33 UTC from IEEE Xplore. Restrictions apply. II. IPMSM MODEL The space-vector diagram of IPMSM is given in Fig. 1. The \ud835\udefc \u2212 \ud835\udefd frame is fixed the stator windings and represents three-phase stator currents in two-axes. The \ud835\udc51 \u2212 \ud835\udc5e frame represents synchronously rotating reference frame in which the \ud835\udc51 \u2212 axis coincides to north pole of the rotor [8]. The mathematical model of IPMSM in the rotating \ud835\udc51 \u2212 \ud835\udc5e frame is given as: [ \ud835\udc63\ud835\udc51 \ud835\udc63\ud835\udc5e ] = [ \ud835\udc45 + \ud835\udc5d\ud835\udc3f\ud835\udc51 \u2212\ud835\udf14\ud835\udc52\ud835\udc3f\ud835\udc5e \ud835\udf14\ud835\udc52\ud835\udc3f\ud835\udc51 \ud835\udc45 + \ud835\udc5d\ud835\udc3f\ud835\udc5e ] [ \ud835\udc56\ud835\udc51 \ud835\udc56\ud835\udc5e ] + [ 0 \ud835\udf14\ud835\udc52\ud835\udf13\ud835\udc43\ud835\udc40 ] (1) where \ud835\udc56\ud835\udc51 and \ud835\udc56\ud835\udc5e are the \ud835\udc51\ud835\udc5e axes armature currents, \ud835\udc63\ud835\udc51 and \ud835\udc63\ud835\udc5e are the \ud835\udc51\ud835\udc5e axes stator voltages, \ud835\udf13\ud835\udc43\ud835\udc40 represents the magnet flux linkage and \ud835\udf14\ud835\udc52 represents the electrical angular velocity of the rotor and \ud835\udc5d is the differential operator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000901_s11012-020-01160-y-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000901_s11012-020-01160-y-Figure1-1.png",
        "caption": "Fig. 1 Circular arc crown gear tooth",
        "texts": [
            " Themain scope of the paper is to develop an analytical method to predict the load sharing ratio and transmission error for crowned spur gears. 2 Load distribution modeling for spur gear pairs with longitudinal modification 2.1 Tooth geometry Gear axial modifications, such as lead crowning, are usually recommended to improve misalignment tolerance and avoid high corner contact at tooth ends, which may otherwise lead to a shortened service life. In this paper a kind of full circular arc crowning, removing slight amounts of tooth from the center on out to reach the edges, is applied on a spur gear pair. As shown in Fig. 1, the heavy lines and thin lines represent the gear tooth flank with and without axial modification. The crowning curve can be described by the following equations R2 \u00bc \u00f0R C0\u00de2 \u00fe \u00f0B=2\u00de2 R2 \u00bc \u00f0R Cp\u00de2 \u00fe z2p ( \u00f01\u00de where R is the radius of circular arc crowning, C0 the amount of axial modification, B the face width and Cp denotes the crowning value of the section where the distance to the tooth center is zp. In Fig. 2, a coordinate system S(x, y) is introduced with setting the symmetrical line of gear tooth as x axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001023_iccre49379.2020.9096444-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001023_iccre49379.2020.9096444-Figure5-1.png",
        "caption": "Figure 5 Motor placements in the model",
        "texts": [
            " Additionally, torque requirements for several critical places were obtained by carrying out motion simulations with the CAD model. VI. CONTROLLING ARCHITECTURE The machine contains overall 9 motors with two high speed brushless motors. Out of these 9 motors 4 motors are to be used in the feeding section. 3 motors are to be used in shuttle positioning system and 2 motors for the ejecting unit. Apart from these motors, 2 optical proximity sensors were used for sensing shuttlecocks and for sensing shots. Figure 5 shows the placement of these motors. VII. DEVELOPMENT OF THE PROTOTYPE To experimentally verify the simulated results and to achieve the multi-shuttle training requirements as described in the paper, a prototype of the machine has been manufactured. Apart from that, the controlling architecture has been implemented and programmed, which was developed in order to regulate the shots and their parametric requirements at a user defined sequence. This proposed machine occupies 2 brushless motors, 2 DC motors, 3 stepper motors and 2 servo motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure6-1.png",
        "caption": "Fig. 6 Dynamic stiffness analysis setup illustration Fig. 7 Baseline cross-member design",
        "texts": [
            " A constant vertical load is applied to the towers for the stress analysis. For the stiffness analysis, an equivalent static stiffness analysis is performed, followed by a dynamic frequency response analysis, to obtain the dynamic stiffness values. Therefore, the degrees of freedom of the system are determined by applying unit loads in translational and rotational conditions. Consequently, the dynamic stiffness is calculated using frequency response analysis. An illustration of the dynamic stiffness analysis setup is shown in Fig.\u00a06. The equation of translational motion of dynamically loaded structures is expressed as follows [4]: (12)K(x)u = F (13)u(x) = K(x)\u22121F. (14)Mu\u0308 + Cu\u0307 + Ku = F 1 3 where M is the mass matrix, C is the viscous damping matrix and K represents the stiffness matrix. u\u0308, u\u0307, and u denote displacement, velocity and acceleration vectors, respectively. F is the applied force vector. Assuming zero initial conditions, one obtains the following harmonic response of a structure by taking the Laplace transform of the transfer matrix as follows: where the complex Laplace transform variable s is substituted by s = j , is the excitation frequency, and j is the imaginary unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.2-1.png",
        "caption": "Fig. 57.2 Approximation of seat base as a plate subjected to bending loads and the equivalent schematic load and support LR = Rider load; LP = Pillion load, RR = Reaction by metal support at the front, RP = Reaction by interfacing part on the rear side",
        "texts": [
            " Ashby [1] highlights the strategy of material selection into four main groups: (a) Translation of the design requirements, (b) screening of the materials using constraints, (c) ranking of the materials using the objective, and (d) seeking the supporting information. In the chosen problem, both strength and stiffness are important parameters for materials selection. Figure 57.1 presents the loads acting on the seat and support (load reaction) points of a seat base in a scooter. In order to simplify the approach to selectmaterials, the seat base can be conveniently considered as a flat panel subjected primarily to bending load (Refer Fig. 57.2), which can further be simplified as a beam bending in 1-D. Thus, the design requirements translate into the parameters as explained in Table 57.1. As thickness of the beam has been considered as the free variable in this case of loading, one could re-arrange Eqs. 57.2 and 57.3 to arrive at the expression for thickness as: where A is the cross-sectional area of the beam, L is the length of the beam, t is the thickness of the beam, \u03c1 is the density of the beam material, F is the load acting on the beam, b is the width of the beam, \u03b4 is the deflection in the beam, C is a constant which depends on the nature of loading (3-point or 4-point), E is Young\u2019s modulus of the beammaterial, and I is the area moment of inertia( I = bt3 12 ) t = ( 12SL3 CEb )1/3 \u3008Stiffness based\u3009 (57"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.4-1.png",
        "caption": "Fig. 34.4 Structural design of fuselage with a boom",
        "texts": [
            " It has a leading edge of 10 mm and a trailing edge of 15 mm. A carbon fibre spar of 8 mm diameter passes through the wing which acts as the main load-bearing element. The fuselage is made of 2 sections, the front part is to house the electronics and the rear part houses the payload. The CG of payload and the CG of aircraft coincide; as a result, the plane can be flown with varying payloads or without payload. The semi-monocoque fuselage design ensures good strength and aerodynamics. As shown in Fig. 34.4, the front section of the fuselage has one bulkhead, three formers and is interconnected by 6 stringers. The second section has 8 formers, 6 stringers, and it contains the wing box. The formers are circular in shape and have a maximum diameter of 172 mm, and the length of the fuselage including the boom is 1210 mm. A boom passes through from the fourth former onwards and is permanently fixed to the fuselage. All formers are made of aeroply, and the stringers are made of balsa. 3 mm thick balsa sheets and aeroply sheets of varying thickness are laser cut into stringers and formers, assembled using a rail and slot mechanism and joined using cyanoacrylate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure15.7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure15.7-1.png",
        "caption": "Fig. 15.7 Fully 3D printed spin coater at different views",
        "texts": [
            " Likewise, packaging foam was cut into cuboids shape of desired dimensions and glued to base of vibration isolation shoe (Fig. 15.6c). All the components are modeled for minimal to no need for support structure for 3D printing allowing good quality of print and minimal wastage of printing material. After 3D printing, some parts may require post-processing like sanding for smoothening surface and drilling for screw fastening. All these factors which was considered for additive manufacturing of parts for spin coater can be easily extended for any part design for products. Final assembled spin coater is as shown in Fig. 15.7. \u2022 Testing of spin coater is done by evaluating two aspects, i.e., speed response for a given user input and vibration characteristics. User interface of the spin coater asks operator for two inputs, i.e., speed constant and spin time. Speed constant is a value ranging from 1050 to 1500; here, each value corresponds to constant RPM rotation of BLDCmotor. Figure 15.8(b) shows good linearity of input speed constantwithmeasuredRPMof theBLDCmotor allowing calibration for required RPM. As shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001889_icem49940.2020.9270999-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001889_icem49940.2020.9270999-Figure8-1.png",
        "caption": "Fig. 8. Field distribution a) Without injection b) Injection",
        "texts": [
            " Cermak, V. Kindl, J. Laksar, Z. Frank, T. Komrska M Authorized licensed use limited to: Western Sydney University. Downloaded on June 15,2021 at 12:33:33 UTC from IEEE Xplore. Restrictions apply. Since the best result has been achieved for #$#%, the implementation of these three components will be used in the study as an example for the induction motor. The simulated nine-phase induction motor was designed according to [1], [7] and [8] The motor has a squirrel cage rotor and its geometry is shown in Fig. 8. The machine parameters are listed in Table III. A full-step winding was used, to achieve the same winding factor for all harmonic components. Fig. 3 shows all the possible connections of the ninephase motor. The three-phase motor can be connected in a star or delta connection. These two connections multiply the voltage magnitudes of any harmonic component by the same ratios '4567 # '89:56 ; . In multiphase machines, different phase shifts occur with higher harmonic components, resulting in a change in the order of the individual waveforms and a different line voltage",
            " Magnetic flux in stator tooth Fig. 7 shows the magnetic flux in stator yoke. Due to harmonic injection, the resulting maximum value is higher by & compared to sine wave. Fig. 7. Magnetic flux in stator yoke Authorized licensed use limited to: Western Sydney University. Downloaded on June 15,2021 at 12:33:33 UTC from IEEE Xplore. Restrictions apply. The presented waveforms were obtained by the star connection. The same results were achieved for the other connections as well. The magnetic flux density distribution in the Fig. 8 shows that there is an increase in stator yoke and decrease in stator teeth. The nine-phase induction motor was coupled with threephase induction motor, which was used as a load (Fig. 9). The motor was measured at constant speed 2 3) . Due to the limited maximum voltage from the nine-phase inverter, the measurement was performed at a reduced voltage, frequency 01 and torque H I The measurement was performed for star and \u201chigher\u201d enneagram stator winding connections. Nonagon and enneagram were not measured, because there are not very suitable for harmonic injection, due to the phase shift described in section IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002244_012013-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002244_012013-Figure1-1.png",
        "caption": "Figure 1. Impeller side room flow.",
        "texts": [
            " Mechanical losses can also include losses of hydraulic braking. But when the pump is running in the designed mode, they are absent. The least researched in that the most interesting are losses on disk friction. Therefore, they will be further considered in this paper. Disk friction losses are understood as the power of friction of the outer surface of the wheels against the liquid. It consists of the friction power of the side surfaces and the friction power of the cylindrical part of the rim. When the disk rotates in a confined space (Fig.1), the liquid between the disk and the wall of the housing rotates at an angular velocity equal to half the angular velocity of the disk. In this case, the moment of friction between the fluid and the disk is balanced by the moment of inhibition due to the friction of the fluid against the walls of the housing. The liquid particles directly adjacent to the surface of the disk rotate at an angular velocity equal to the speed of the disk. The centrifugal forces acting on them are not balanced by the pressures in the main flow"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001609_0021998320960531-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001609_0021998320960531-Figure19-1.png",
        "caption": "Figure 19. Out of plane compression behaviour of sandwich panels with 30 (left) and 45 (right) corrugated core.",
        "texts": [
            " The load was applied as displacement on the reference point that was coupled with the upper surface of upper face sheet. Similar approach was also used by Bing et al.34 to investigated the flatwise compression properties of glass fibre reinforced PP corrugated sandwich panels. Rest of the details about the numerical model are discussed in the Bending stiffness of SrPPSBs section. The deflection behaviour of SrPP corrugated structures at different stages under flatwise compression is shown in Figure 20 and Figure 19 along with the mid span deflection d at different points. The first mode buckling is shown by brown arrows, second mode is shown by red arrows while contact of struts with upper FS is shown by white arrows. The load displacement curves for corrugated sandwich structures are shown in Figure 21(a) while more magnified curves with respect to deflection are shown in Figure 21(b). Peak load values are shown in Figure 21(c). The 30 corrugated sandwich structures first show linear behaviour up to d\u00bc 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000440_s11071-019-05457-w-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000440_s11071-019-05457-w-Figure4-1.png",
        "caption": "Fig. 4 Phase-space diagrams from simulation; x-axis is the voltage across R, i.e., VR in volt, and y-axis is the voltage across the capacitor C1 in volt. The parameters are R = 156.77 , L = 100 mH, C1 = 1.629 nF, C2 = 56.2 pF and driving frequency f = 10 kHz. a Period-1 orbit at Vamp = 1.0 V, b narrow-band chaos at Vamp = 1.075 V, and c period-2 orbit at Vamp = 1.3 V",
        "texts": [
            " The borderline between the two compartments is represented by the condition that the value of q reaches qref exactly at the Poincar\u00e9 observation instants (maximumpositive amplitude of the input sine wave). Figure 3 shows a few numerically obtained time-series waveforms for the circuit. When the capacitor voltage (VC1) does not touch the reference voltage (Vref ), the comparator output is 1, i.e., high. This is the condition before the border collision, and the system shows a period-1 attractor. Figure 4a shows the phase-space diagram of the period-1 attractor. But when the VC1 just touches Vref , i.e., at the bifurcation point, a chaotic attractor develops (Fig. 3b). The corresponding phasespace diagram is shown in Fig. 4b. The chaotic attractor exists for a narrow range of the parameter Vamp. As Vamp increases, a period-2 orbit develops. It has two loops; in one loop, the capacitor voltage crosses Vref , and in the other loop it does not. The corresponding phase-space diagram is shown in Fig. 4c. Figure 5 shows the numerically obtained bifurcation diagrams of the system. In the figure, Vamp and Vref are taken as the bifurcation parameters, respectively, keeping the rest of the parameters fixed. It exhibits a border collision bifurcation from a period-1 orbit to a period-2 orbit through a large amplitude chaotic oscillation at the bifurcation point. The same behavior has been reported in the analogous mechanical impacting system [32]. Figure 6 shows the attractor at the grazing condition in the discrete-time phase space"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure17-1.png",
        "caption": "Figure 17. Drawing of the interior frame of the car door.",
        "texts": [],
        "surrounding_texts": [
            "The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article."
        ]
    },
    {
        "image_filename": "designv11_63_0001673_icuas48674.2020.9214061-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001673_icuas48674.2020.9214061-Figure4-1.png",
        "caption": "Fig. 4. Distances of the forces acting on the UAV.",
        "texts": [
            " FB = [ FB x FB y FB z ]T and \u0393B = [ \u0393B x \u0393B y \u0393B z ]T FB x = c\u00b5 (FR1 +FR2)\u2212 c\u00b5 (DW1 +DW2 +DWe +DWr) \u2212s\u00b5 (LW1 +LW2 +LWe) FB y = LWr FB z = s\u00b5 (FR1 +FR2)\u2212 s\u00b5 (DW1 +DW2 +DWe +DWr) +c\u00b5 (LW1 +LW2 +LWe) (6) \u0393B c =  s\u00b5 (YR1FR1\u2212YR2FR2)+ c\u00b5 (YW1LW1\u2212YW2LW2) \u2212Ze (c\u00b5DWe + s\u00b5LWe)+Xe (\u2212s\u00b5DWe + c\u00b5LWe) c\u00b5 (YR2FR2\u2212YR1FR1)+ s\u00b5 (YW1LW1\u2212YW2LW2) \u2212s\u00b5 (YW1DW1\u2212YW2DW2) \u2212Xr (s\u00b5DWr) +c\u00b5 (YW1DW1\u2212YW2DW2)\u2212XrLWr  (7) \u0393B g =  s\u00b5Izrq(\u03c9R2\u2212\u03c9R1) c\u00b5Izrr (\u03c9R2\u2212\u03c9R1)+ s\u00b5Izr p(\u03c9R1\u2212\u03c9R2) c\u00b5Izrq(\u03c9R1\u2212\u03c9R2)  (8) \u0393B D = s\u00b5 (YR2DR2\u2212YR1DR1) 0 c\u00b5 (YR1DR1\u2212YR2DR2)  (9) Where L and D are the aerodynamic lift and drag respectively generated by the wings (W), the horizontal (We) and vertical (Wr) stabilizer. (\u0393B c ) the main moments provided by the control surfaces, (\u0393B g ) the gyroscopic moment and (\u0393B D) the drag moment due to the propeller drag force. With YR, YW , Xe, Xr, Ze as the distances of the forces acting on the motors, wings, horizontal and vertical stabilizer respectively (Figure 4). For the transition flight mode, the complete mathematical model is simplified according to the following aerodynamic considerations. (1) The tilt angle is from \u00b5 = 0\u25e6 to \u00b5 = 90\u25e6. (2) The \u03c6 and \u03c8 angles of the rotational dynamics are small, such that, cos(a) = 1 and sen(a) = a. (3) The forces in y direction are small compared to x and z. (4) Only the main moments are considered. Therefore, the vehicle dynamics can be written as: mz\u0308 = FB x (s\u03b8 +\u03c6\u03c8)+FB z (c\u03b8)\u2212mg \u03c6\u0308 = 1 c\u03b8 [ \u2212\u03b8\u0308\u03c8 + \u03c6\u0307 \u03b8\u0307s\u03b8 + \u03c6\u0307 \u03c8\u0307c\u03b8\u03c8\u2212 \u03b8\u0307 \u03c8\u0307 + 1 Ix( \u0393B x \u2212 ( \u2212\u03c6\u0307c\u03b8\u03c8 + \u03b8\u0307 )( \u03c6\u0307s\u03b8 + \u03c8\u0307 ) (Iy\u2212 Iz) )] \u03b8\u0308 = \u03c6\u0308c\u03b8\u03c8\u2212 \u03c6\u0307 \u03b8\u0307s\u03b8\u03c8 + \u03c6\u0307 \u03c8\u0307c\u03b8 + \u03b8\u0307 \u03c8\u0307\u03c8 + 1 Iy( \u0393B y \u2212 ( \u03c6\u0307c\u03b8 + \u03b8\u0307\u03c8 )( \u03c6\u0307s\u03b8 + \u03c8\u0307 ) (Iz\u2212 Ix) ) \u03c8\u0308 =\u2212\u03c6\u0308s\u03b8 \u2212 \u03c6\u0307 \u03b8\u0307c\u03b8 + 1 Iz ( \u0393B z \u2212 ( \u03c6\u0307c\u03b8 + \u03b8\u0307s\u03c8 )( \u2212\u03c6\u0307c\u03b8\u03c8 + \u03b8\u0307 ) (Ix\u2212 Iy) ) (10) with the following forces and moments, FB x = c\u00b5 (FR1 +FR2)\u2212 c\u00b5 (DW1 +DW2 +DWe +DWr) \u2212s\u00b5 (LW1 +LW2 +LWe) FB z = s\u00b5 (FR1 +FR2)\u2212 s\u00b5 (DW1 +DW2 +DWe +DWr) +c\u00b5 (LW1 +LW2 +LWe) \u0393B x = s\u00b5 (YR1FR1\u2212YR2FR2)+ c\u00b5 (YW1LW1\u2212YW2LW2) \u2212s\u00b5 (YW1DW1\u2212YW2DW2) \u0393B y =\u2212Ze (c\u00b5DWe + s\u00b5LWe)+Xe (\u2212s\u00b5DWe + c\u00b5LWe) \u2212Xr (s\u00b5DWr) \u0393B z = s\u00b5 (YW1LW1\u2212YW2LW2)+ c\u00b5 (YW1DW1\u2212YW2DW2) \u2212XrLWr (11) As mentioned in section I, investigations based on the study of transition dynamics generates a reference trajectory (Figure 5), which consists of a sequence of equilibrium points that correspond to specific wing-tilt angle and cruise forward speed combinations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000112_1077546319874921-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000112_1077546319874921-Figure2-1.png",
        "caption": "Figure 2. Structure of the semi-active hydraulic damping strut.",
        "texts": [
            " The semi-active HDS is a kind of hydraulic damper that can achieve different dampings by opening or closing the outer channel, which is controlled by the vehicle ECU (Electronic Control Unit). As the vehicle is in the process of in situ shift, the shift signal is transmitted to the vehicle ECU through the CAN-BUS (Controller Area Network-Bus), then the ECU sends the instruction to make the solenoid valve of the semi-active HDS in working condition to close the outer channel, and a larger damping is produced in this process. The outer channel will be opened if the shifting is finished, and then the lower damping is produced. The structure of the semi-active HDS is shown in Figure 2. The design content of the semi-active HDS mainly includes the outer channel diameter, the damping hole diameter, and the quantity of the inner piston. The outer channel diameter of the semi-active HDS has a great effect on the vibration isolation performance of the powertrain mounting system, for example the engine is at an idle speed and high-speed cruise state, and it can be analyzed according to the design method of the powertrain mounting system. The inner damping hole diameter and quantity of the piston have a great effect on the vehicle transient shock and vibration in the process of in situ shift, and this paper mainly analyzes the optimization method of the damping hole diameter of the inner piston"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002017_cce50788.2020.9299157-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002017_cce50788.2020.9299157-Figure7-1.png",
        "caption": "Fig. 7. Pressure Contours and Flow-Path Lines (Re = 0.216e6, \u03b1 = 14o).",
        "texts": [
            " In the case of the wing and the airframe, the slope is positive, this may be due to the fact that neither the selected airfoil nor the fuselage geometry are appropriate for turbulent flow. To compensate for this instability, the components of the airplane could be easily changed so that the center of gravity lies in a position that overcomes the nonrestoring pitching moment. The pressure variation over the surface with flow-paths at 14o, where maximum lift coefficient was achieved, is presented in Fig. 7a. Sanity checks were done and it was verified that the pressure was higher at the lower surface, which indicates a resultant lift force. Meanwhile, Fig. 7b shows flow-paths near the wing tips, where tip vortices and consequently more induced drag are being generated. Induced drag is an unfavorable three-dimensional effect [10], an effect that could be reduced by adding winglets on the airplane. The effect of the motor pods in the structure is also noticeable, when the flow collides with them several disturbance and more drag is being generated. Since the prototype concept was directly modeled with no changes in geometry future work may include adding winglets and observing the flowpaths again"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000913_jfm.2020.272-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000913_jfm.2020.272-Figure12-1.png",
        "caption": "FIGURE 12. Initial configuration of crossing interactions, during which there is an offset of 0.5a compared to the side-by-side interaction. Here, \u03b4l is the minimum separation distance between the membranes of two swimmers.",
        "texts": [
            " This is because larger \u03b8 ini 12 generates higher pressure between the two near-contact surfaces, which results in large membrane tension, even though the trajectories were not strongly affected by the pressure. Figure 11(a) shows the effect of deformability, Ca, on the relative orientation of cells (\u03b8 ini 12 = \u03c0/3). The final angle slightly decreases as Ca is increased. The decrease in scattering angle may cause a decrease in the self-diffusivity of cells in the suspension. Isotropic tension in the anterior region, \u03c4 ant 0 , is also influenced by Ca. Note that \u03c4 ant 0 decreases as Ca is increased (figure 11b). 3.3. Crossing interaction Figure 12 shows the initial configuration of crossing interactions. Swimmer 1 is initially oriented in the x-direction, whereas swimmer 2 has a relative orientation angle of \u03b8 ini 12 . To ensure their paths cross, we use an offset of 0.5a compared to the side-by-side interaction. In this case, the posterior region of swimmer 1 and the anterior region of swimmer 2 come into close contact. The trajectories of a swimmer\u2019s volumetric centres with \u03b8 ini 12 = \u03c0/6, \u03c0/3, \u03c0/2 and 2\u03c0/3 are shown in figure 13 (Ca= 0",
            " The rate at which an escape reaction happened during crossing interactions was investigated in Ishikawa & Hota (2006). It was found that approximately 20 % of Paramecium that were ahead showed an escape reaction regardless of the initial angle \u03b8 ini 12 , and the maximum rate of escape reaction occurred when \u03b8 ini 12 =\u03c0/2. To reproduce the experimental settings, we performed a simulation of crossing interactions with varying \u03b8 ini 12 . To focus on the effect of initial angle \u03b8 ini 12 , we adjusted the offset of swimmer 2 (figure 12) to achieve a minimum separation distance between the two membranes of the two swimmers during the interaction in the range \u03b4l = 0.069a \u223c 0.07a, regardless of \u03b8 ini 12 . The results of isotropic tension, \u03c40, of swimmer 1 are shown in figure 18(a). The tension in the posterior region \u03c4 pos 0 of the swimmer that is ahead is considerably increased regardless of \u03b8 ini 12 . This result is consistent with the experimental observations, in which approximately 20 % of Paramecium that were ahead had an escape reaction regardless of the initial angle \u03b8 ini 12 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000185_1.5131964-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000185_1.5131964-Figure2-1.png",
        "caption": "FIGURE 2. The appearance of modular manipulators of the EBM machine",
        "texts": [],
        "surrounding_texts": [
            "The starting materials were following: welding wire made of Ti-6Al-4V a diameter of 1.2 mm and stainless steel AISI 308LSi with a diameter of 1.2 mm. The chemical composition of the materials is presented respectively in Tables 1 and 2. In 2015 a modular installation for electron-beam fusion of powders and cladding with wire was created and have been constantly modernized at TPU from then on. The setup includes a vacuum chamber with an electron-beam gun with a plasma emitter and modular manipulators, enabling layer-by-layer cladding of powders by electron-beam melting (EBM) or dimensional welding with wire. The appearance of the setup and the structural schemes are shown in Figures 1\u20133. 020097-2 The technical characteristics of the machine are as follows: accelerating voltage is 40 kV, base pressure is 5\u00d710-3 Pa, maximum beam current is 200 mA, minimum current beam diameter is 150 \u03bcm, build area dimensions are 150\u00d7150 mm, power consumption is 6 kW. The electron beam is generated by a source with plasma emitter. A distinctive feature of the source is the use of gas discharge plasma as the electron emitter. Such design provides a number of new operational benefits in comparison with thermal cathode sources. For instance, plasma-emitter sources preserve performance at elevated operating pressure. In addition, they demonstrate long life under the action of metal vapors, including refractory ones, and gas emissions from the melting zone. They are successfully used in traditional electron-beam technologies such as welding, surfacing. The machine is intended for manufacturing parts with complex geometries from metal powders or a metal wire. The included software enables modular exchange and synchronized control of all installation components according to the task using digital G-codes. Initially, the Linux NC system was designed to control turning, milling machines, 3D printers and other mainly machine-building equipment. However, the potential laid by the developers in the system makes it possible to develop a more complex system, for example, an electron-beam installation. During the work, samples of Ti-6Al-4V titanium alloy and AISI 308 steel were printed on an electron-beam 3Dprinter at an accelerating voltage of 30 kV and a beam current of 15 to 20 mA (depending on the distance from the substrate). The input power varied from 450 to 600 watts. The focused beam (150 \u03bcm in diameter) was performing circular scan with 4 mm in diameter. The wire was fed to the beam area, and the sample geometry was achieved by moving the table along three axes. The distance between the tracks (hatch distance) was 4 mm, the layer height was 0.8 mm, the beam was moving in in the horizontal plane in zigzag pattern. The samples shown in Figures 4 and 5 demonstrate the installation capabilities for obtaining both bulk and square samples with wall thickness in one pass. The obtained samples allow studying the structure and microstructure of the alloys depending on the energy, geometric and kinetic parameters of the beam and the geometry of the obtained samples. To assess the porosity of the specimens and to perform mechanical testing, the samples were cut into rectangular specimens shown in Figure 6 by the electroerosion method. 020097-3 Vickers hardness was determined on the transverse sections of the specimens using an Tukon 2500 Automated Knoop/Vickers Hardness Tester (2500-6 modification, Instron, USA) with a load of 1 kg. The indentation time was 10 seconds. \u0425-ray computed tomography (XCT) studies were performed using a scanner OREL-MT developed in TPU. The scanner equipped with XWT-160-TC X-ray tube (X-ray WorX, Germany) and PaxScan-2520V flat panel X-ray detector (Varian, the USA). The scanning parameters were as follows: accelerating voltage \u2013 130 kV, copper filter of 0.2 mm thickness, angle step of 0.3 degrees, 1200 shadow projections, voxel size \u2013 10 um."
        ]
    },
    {
        "image_filename": "designv11_63_0002173_s40313-020-00669-7-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002173_s40313-020-00669-7-Figure2-1.png",
        "caption": "Fig. 2 Coordinates system for the state variables",
        "texts": [
            " The navigation of the vehicle is given by a dual GNSS/INS system to ensure the accuracy in the position, velocity and accelerationmeasures. The communication system is based on RF transceivers, with a range of 60 miles (96.56 km) in the 900 MHz spectrum. The control architecture is distributed between Texas Instrument launchpad microcontrollers and an NVIDIA Jetson TX-2 central computer running on GNU/Linux operative system and powered by a robot operating system (ROS). Themathematical model of a vessel can be expressed by considering its horizontal motions and following the SNAME (1950) notation. Figure 2 presents the inertial-frame coordinates of the vehicle composed of surge position x , sway position y and yaw position \u03c8 . The body-frame coordinates of the vehicle are composedof surge velocityu, swayvelocity v and yaw velocity r . According to the matrix representation in three degrees of freedom (3-DOF), the position vector relative to inertial-frame is represented by \u03b7 = [ x y \u03c8 ]T , the velocity vector relative to the body-frame is represented by \u03bd = [ u v r ]T , and the vector of the generalized input forces is represented by \u03c4 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002315_j.matpr.2020.12.487-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002315_j.matpr.2020.12.487-Figure10-1.png",
        "caption": "Fig. 10. Process of developing IWP pattern insole.",
        "texts": [
            " In all the patterns, region of stress is concentrated at every unit cell except for IWP pattern where the maximum stress is concentrated at the last layer of unit cell. This indicates high compression strength which is known property of IWP pattern. Depending on this criteria and results obtained above, IWP pattern is chosen as infill for insole. Along the total path of insole IWP patterns are embedded and partitioned with enclosed surface of insole. Further the external layer of insole is added with 1 mm thickness and merged with internal IWP patterns. The process of generating IWP insole is shown in Fig. 10. Optimized results with respect to weight for designed, modified and IWP pattern insole are shown in Table 5. Fig.8. (A) 3D tetra mesh; (b) mesh quality parameters, loads and boundary conditions. Flat foot is one of the major cause for several foot problems and lower limbalignment.Neglectingthiswouldaffectbodypartsparticularlyalongsagittalplane.Pesplanusoccurswhenplantarfascia ligament falls down from normal position. To overcome this problem, patientspecific insole isdesignedandoptimizedfor improvingalignmentofbodyandprovidingbettercomfortevenduringregularusage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure34.8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure34.8-1.png",
        "caption": "Fig. 34.8 Streamlines of the aircraft",
        "texts": [
            " Since the flow is incompressible and the analysis is performed at a lowmach number, to reduce the computational cost a pressure based solver is selected. The result of the analysis provided a lift of 83.955 N and a drag of 5.655 N. The initial settings and fluent inputs of this analysis are similar to the analysis of the wing. The model had a skewness of 0.9, and an inflation layer was given to the wing with y + 1. The result of the analysis provided lift of 74.38 N and a drag of 8.42 N. From the streamlines, we were able to tell that the tail was not falling under the wake region of the wing (Fig. 34.8). Parameter Values Overall length 220 mm Overall width 140 mm Overall height 100 mm Maximum capacity 2.2 l Time to empty the full tank 2 min 56 s Area sprayed in 2 min 56 s 4941 m2 (1.22 acres) Payload properties of designed chemical (pesticide) spraying tank are shown in Table 34.11. The payload can be easily swapped; other payloads which can be used are crop-monitoring equipment, artificial pollination set-up, land surveying and mapping equipment. Aircraft can carry maximum of 4 kg of payload, while maintaining sufficient factor of safety for the aircraft to fly at its peak performance, which meets the design requirement as stated in Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002279_1.4006737-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002279_1.4006737-Figure3-1.png",
        "caption": "Fig. 3 Overview of the first building inner facility model",
        "texts": [
            " The proposed method can be effectively used for purposes such as mass-evacuation planning and evacuee action prediction. 2.1 Simulation Platform and Evacuation Route. Figure 1 shows a 3D model of the Tokyo Tech High School of Science and Technology (TTHS) as a simulation platform used to compare the evacuation drill results discussed later herein. SilTools [8] was used for the 3D modeling. Figure 2 shows the school inner facility model with its global coordinate Rg, where through are possible evacuation routes. Each evacuee proceeds to the playground through adequate stairways and ground routes. Figure 3 shows an overview of first building inner facility and its local coordinate frame. Students waiting for the drill to begin are also located in the building. The students in the three southside classrooms proceed outside through the nearest corridor or stairway. Similarly, students on the second and fourth floors on the north side of the building proceed down the north inside stairway through the nearest corridors, whereas students on the third floor proceed down the north outside stairway. A similar scenario is applied to the evacuees in other buildings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002194_s12206-020-1201-5-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002194_s12206-020-1201-5-Figure5-1.png",
        "caption": "Fig. 5. The test rig of planetary gearbox.",
        "texts": [
            " This section mainly verifies the alpha stable distribution characteristics of planetary gearbox condition monitoring signals, extracts the alpha stable distribution parameters of planetary gearbox performance degradation process, and gives the new method of degradation feature extraction. For ease of description, the planetary gearbox degradation experiment and condition monitoring data collection methods are introduced first. Planetary gearbox degradation data is derived from a lifecycle degradation experiment. The planetary gearbox test rig is shown in Fig. 5. The test rig consists of drive motor, magnetic powder brake, speed and torque sensor, test gearbox and other main parts. The test gearbox is a single-stage NGW-11 planetary gearbox with a transmission ratio of 12.5, the specific structural parameters of test gearbox are shown in Fig. 6. During the experiment, four vibration acceleration sensors are arranged in the planetary gearbox case, and the specific mounted location of sensors are shown in Fig. 7. Sensors 1# and 2# are installed at the input shaft end, sensor 1# is horizontal, sensor 2# is axial direction, sensor 3# is installed at the top of the gearbox, sensor 4# is installed at the output shaft end and is axial direction, and all sensors are installed near the ring gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002524_tmag.2021.3065901-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002524_tmag.2021.3065901-Figure4-1.png",
        "caption": "Fig. 4. Geometry, magnetic flux lines and torque exerted by the switched reluctance machine for four different skew angles \u03b3skew \u2208 {0\u25e6, 10\u25e6, 20\u25e6, 30\u25e6}.",
        "texts": [
            " A remedy would involve the introduction of a helicoidal coordinate system [22] in the rotor slices. A further modeling improvement could be the application of a quasi-3-D FE discretization [23]. However, these measures would further complicate the overall multi-slice 2-D solver. Even if the gap between such a highly specialized solver and a standard 3-D solver becomes smaller, it will never attain the accuracy of a full 3-D approach. The 2-D FE formulation with skew is illustrated for a single-slice 2-D model of a switched reluctance machine. The torque when exciting one coil is shown in Fig. 4 for four different skew angles. The results clearly show that the single-slice model is already capable of dealing with skew. Up until now, this feature was only given by the substantially more difficult approach proposed in [8]. A single slice, however, cannot account for the axial dependence of the stator and rotor fields. Especially, for induction machines, the axial variation of the ferromagnetic saturation is significant. The skew interface condition technique proposed here can be applied in a multi-slice setting as well"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure4-1.png",
        "caption": "Fig. 4 Topology and topography optimisation",
        "texts": [
            " The relation between the elastic modulus and the material density at point x is expressed by following equation [1, 2]: where E0 is the elastic modulus of the fully solid material. The volume of an element is obtained by the following equation [1, 2]: where Vi is the volume of the ith element and Vi 0 is the original volume of the ith element. For an isotropic material, a non-zero base plate thickness can be defined. However, the base plate thickness must be zero for a composite plate or a plate with anisotropic material properties, which can be seen as a limitation of the current development. The Topology and Topography Optimisation is shown in Fig.\u00a04. For the incorporation of topology and topography optimisation into the design, the gauge optimisation to identify the optimal part gauges and minor adjustments to the resulting design to ensure performance were carried out. (6)D = pE 1 \u2212 2 \u23a1 \u23a2\u23a2\u23a3 1 0 1 0 0 0 1\u2212 2 \u23a4 \u23a5\u23a5\u23a6 (7)K(x) = n\u2211 e=1 xp e K0 e (8)Ei = Ei ( xi ) = x p i E0, xi \u2208 [0, 1] (9)Ei = Ei ( xi ) = Emin + x p i ( E0 \u2212 Emin ) , xi \u2208 [0, 1] (10)E(x) = f ( x)E0 (11)Vi = \u222bV0 i (x)dV 1 3 Static stiffness is calculated by applying unit loads to the engine mounting positions in all degrees of freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001555_physreve.102.032701-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001555_physreve.102.032701-Figure2-1.png",
        "caption": "FIG. 2. Schematic diagram of the meniscus region of a free standing smectic film. Only the left section of a film taken between two parallel walls (shaded region) is shown. The thickness h of the meniscus decreases as one moves away from the wall along X, reaching a flat part with thickness hF at X = Xmax. Edge dislocations located at the midplane of the meniscus mediate the variation in h. The Burgers vector of the dislocation (shown by the symbol T rotated by 90\u00b0) is equal to the layer thickness d near the thinner parts and several times d in the thicker parts. The surfaces of the meniscus have a circular shape with a radius RM \u223c 100 \u03bcm. As the sample is cooled towards SmA to SmC transition temperature (TAC), the director in the layers near the surface develops a tilt angle, and the corresponding c vector can be expected to develop a bend distortion by following the surface profile, shown by the dashed lines near the surfaces. It is argued in the text that this actually costs a positive energy, which is reduced by the formation of stripe distortions of the thinner parts of the structure with periodicity along Y. is the small width of a slice with thickness h used in the theoretical analysis.",
        "texts": [
            " The sample can be annealed so that the central part of FSFs has a flat structure with a well-defined number of layers, and investigations on such layers have led to various discoveries like that of hexatic phases, etc. The flat central part is connected to the walls at the edge of the hole through a meniscus region in which the thickness increases continuously, mediated by edge dislocations in the layer structure. The dislocations are repelled by the large surface tension of smectic LCs (\u03b3 \u223c 0.02 N/m) to the center of the meniscus (Fig. 2). Detailed experimental and theoretical studies [2] on SmA menisci have shown that in the thinner parts close to the center, the dislocations have a Burgers vector equal to the layer spacing d . Closer to the wall, large Burgers vectors ( 20 or so) mediate in the thicker parts, which gives rise to a two-dimensionally deformed structure made of parabolic focal conics. The meniscus surface has a circular shape, with a radius RM 100 \u03bcm, except very close to the central flat region, in which it has an essentially linear slope [3]",
            " If the sign of k is reversed, that of c is reversed as well, and only those curvature distortions in which the c- and k-dependent terms totally add up to an even number are allowed. Typically, the SmC stripes of a given periodicity form a closed chain without coming into contact with any solid boundaries, in the meniscus region around the central flat film. Similarly, the coronal stripes around a deposited particle on the central part of the FSF also form a closed chain. This allows us to simplify the problem, and assume the FSF of SmA LC to be formed between two parallel vertical walls (Fig. 2) with edges lying in the YZ plane, and extended along the Y axis, so that the equilibrium SD structure can be found by minimizing the energy density. Both the top and bottom surfaces of the meniscus have a circular shape in the XZ plane, with a radius RM , except very close to the boundary with the central film of uniform thickness. The surface layers with k pointed towards the center of the circle develop a tilt angle even a few degrees above TAC. We first consider the temperature of the sample to be below TAC, so that it is in the SmC phase",
            " The bend vector itself has a splay distortion as it is aligned with k. This of course would lead to a crowding of the ends of the tilted molecules at the surface, which is not favored. In analogy with the saddle-splay elasticity in nematics, we introduce the elastic term kSS \u2207 \u00b7 (c\u2207 \u00b7 c + c \u00d7 \u2207 \u00d7 c) which is of course a surface term. The above argument shows that kSS is negative, resulting in a positive energy density in the medium. Considering a thin vertical slice of width , across which the meniscus height h can be assumed to be constant (Fig. 2), the average energy density of the slice due to the above term is given by FSS = \u22122kSS/(hRM ). (1) The factor 2 in the above arises as there are two equivalent surfaces in the meniscus. A surface c field which is perpendicular to the plane of the paper in Fig. 2 does not necessarily reduce the energy density, as the ends of the tilted molecules 032701-3 are still crowded at the surface, costing energy. Introducing the vector b = k \u00d7 c, it is clear that b will have the bend distortion in the latter case. We assume that kSS introduced above adequately captures the physical process, and ignore b-dependent terms in the following. We can note that around layer steps in the FSF the c vector aligns parallel to the step edge, i.e., the dislocation [3,17]. However, in the meniscus the edge dislocation is located at the center, in which the c vector has been found to have the orthogonal orientation, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002378_tasc.2021.3059390-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002378_tasc.2021.3059390-Figure2-1.png",
        "caption": "Fig. 2. 3-D finite element model of a superconducting LSM unit and center line of racetrack coil.",
        "texts": [
            " Therefore, the 2-D finite element method is challenged by the arc parts of racetrack coils and the 3-D finite element method is challenged by the large geometry size of superconducting LSM. Considering the calculating time and reliability, a real-scale 3-D finite element model with a position-adjustable secondary is employed and established according to the configurations in Table I. The model has been verified in our previous work [10]. In the 3-D finite element model of superconducting LSM shown in Fig. 2, the electromagnetic force F is calculated based on the Lorentz force equation F = \u222e [(iyBz \u2212 izBy) i\u2212 (ixBz \u2212 izBx) j +(ixBy \u2212 iyBx)k] dl, (1) i, j, k are the unit vectors in the x-, y- and z-direction; (ix, iy, iz) is the vector of current in the primary coil; (Bx, By, Bz) is the vector of magnetic flux density produced by the secondary coils. Furthermore, the lengths of air region in y-direction and z-direction should be long enough for the free dissipation of exciting magnetic field produced by superconducting coils",
            " Similarly, the effects of z-direction offset on those forces are investigated. As mentioned above, the 3-D relative position between the primary and secondary of superconducting LSM varies with the operating of EDS train, which is significantly different form the traditional applications of motors. Thus, the electromagnetic characteristics of superconducting LSM should be studied at the rated condition firstly. The geometry structure and rated condition of superconducting LSM are designed and shown in Table I and Fig. 2. In Fig. 1, the gap and dy are 257 mm and 0 mm at the rated working condition of superconducting LSM, respectively. The xoy planes with z = 0 mm and z = 257 mm are defined as the central plane of the primary windings and the secondary windings, respectively. The central points (x,y) of the first coil of the primary and the first coil of the secondary are defined as (0 mm, 0 mm) in the 3-D finite element model. Fig. 3 shows the contour plot of Bz (the z component of no-load magnetic field) in the central plane of the primary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure7-1.png",
        "caption": "Figure 7. Third and fourth eigenmode (95.8Hz and 98.1Hz)",
        "texts": [],
        "surrounding_texts": [
            "Composite materials, finite element analysis, automotive engineering, vibrational behavior, design optimization Date received: 21 March 2021; accepted: 5 May 2021"
        ]
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure49.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure49.3-1.png",
        "caption": "Fig. 49.3 Placement of components inside the enclosure",
        "texts": [
            " The functional requirements were derived by discussions with expert ophthalmologists and are as follows: \u2022 Usable by general medical practitioners \u2022 Provide accurate indication of normal or elevated IOP \u2022 User-friendly and not require complex instructions and training \u2022 Should not require aesthetic drops (not permitted for home use) \u2022 Minimize patient discomfort \u2022 Cost effective. The proposed design is in the form of headgear with a screening module attached to it. The patient can directly wear this headgear, increasing the feasibility and convenience for diagnosis for the clinician. This overcomes the difficulty in keeping the device stationary above the patient\u2019s eye during applanation. The screening module houses five parts in an enclosure\u2014screw cap, force sensitive resistor, spring and indenter. (Fig. 49.3) For assessment of IOP, the force resistive sensor (FSR) is placed on the indenter (Fig. 49.3). Using the Imbert-Fick principle, the force on the sensor can be obtained by multiplying the area of the indenter tip by the pressure of the eye. An additional spring force also acts on the sensor. The function of the spring is simply for pulling the indenter back to its original position after it is pushed for applanation. The total force on the sensor would be checked, and it should be less than the limiting force. If the IOP of the patient crosses the limiting value, then the increased load on the FSR will signal through a red LED"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001470_2050-7038.12588-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001470_2050-7038.12588-Figure17-1.png",
        "caption": "FIGURE 17 Magnetic flux density distribution",
        "texts": [
            " By applying notch to the stator, no change in the total magnet amount will occur and the negative situation mentioned in the previous section will be eliminated. In this method, the notch must be deep enough. Otherwise, the notches opened will not behave similar to the stator slot and may be ineffective in reducing cogging torque.49 The change in the flux density of the initial design and this design in the airgap is given in Figure 16. In the notch parts opening to the stator, the flux density decreases as the reluctance increases. The distribution of the flux density obtained on the stator as a result of magnetostatic analysis is presented in Figure 17. When the core material used is considered, it is seen that the flux density remains within the limits. FIGURE 13 Air-gap flux density: A, radial and tangential components; B, for a pole pitch FIGURE 15 Cogging torque variation according to the rotor position In permanent magnet machines, when each magnet passes through the stator teeth, the reluctance seen by the magnet under the slot opening changes due to the length of the flux path. Hence, the slot opening creates a varying reluctance for the flux produced by the magnet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001706_s12555-020-0031-7-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001706_s12555-020-0031-7-Figure5-1.png",
        "caption": "Fig. 5. 3R planar manipulator, a three-link manipulator with revolute joints.",
        "texts": [
            " We may generalize the observation made for a 2R manipulator in section 3.1 for a N degrees-of-freedom planar manipulator. The term \u03c4i = \u03a3N i \u03c4\u0303i is the total interactive moment between the links i\u22121 and i about Zi, while \u03c4\u0303i is the interactive moment between the links i\u22121 and i about Zi if the links (i+1)\u2212N do not exert any moment on the ith link. Using the special structure of the dynamics of planar manipulators discussed in the previous section, we now develop a distributed cooperative control scheme for a 3R planar manipulator as illustrated in Fig. 5. For a 3R planar manipulator, the dynamics ( (1)) maybe now re-written as \u03c43 = \u03c4\u03033; \u03c42 = \u03c4\u03033 + \u03c4\u03032; \u03c41 = \u03c4\u03033 + \u03c4\u03032 + \u03c4\u03031 (13) with \u03c4\u0303i = M\u0303i(\u03b8)\u03b8\u0308 +V\u0303i(\u03b8 , \u03b8\u0307)+ G\u0303i(\u03b8). (14) Here, M\u0303i(\u03b8) is the ith row of M\u0303(\u03b8) and V\u0303i(\u03b8 , \u03b8\u0307) is the ith element of the vector V\u0303 (\u03b8 , \u03b8\u0307), and G\u0303i is the ith element of the vector G\u0303(\u03b8) The joint-wise dynamics of the 3R planar manipulator written in standard form as given in (2) is \u03c4i = Mi(\u03b8)\u03b8\u0308 +Vi(\u03b8 , \u03b8\u0307)+Gi(\u03b8). (15) Comparing (15) with (13) and (14) we have M1(\u00b7) = M\u03031(\u00b7), M2(\u00b7) = M\u03031(\u00b7)+ M\u03032(\u00b7), M3(\u00b7) = M\u03031(\u00b7)+ M\u03032(\u00b7)+ M\u03033(\u00b7), (16) V1(\u00b7, \u00b7) = V\u03031(\u00b7, \u00b7), V2(\u00b7, \u00b7) = V\u03031(\u00b7, \u00b7)+V\u03032(\u00b7, \u00b7), V3(\u00b7, \u00b7) = V\u03031(\u00b7, \u00b7)+V\u03032(\u00b7, \u00b7)+V\u03033(\u00b7, \u00b7), (17) G1(\u00b7) = G\u03031(\u00b7), G2(\u00b7) = G\u03031(\u00b7)+ G\u03032(\u00b7), G3(\u00b7) = G\u03031(\u00b7)+ G\u03032(\u00b7)+ G\u03033(\u00b7)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure6.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure6.1-1.png",
        "caption": "Fig. 6.1 Rotating phasor with variable length and speed of rotation",
        "texts": [
            " The DHB allows one to find equations of a transient oscillatory motion in a vicinity of the periodic solution given by (6.5). Consequently, analysis of these transient oscillatory motions leads to the conditions of orbital stability. The DHB principle was formulated in [15]. Let us briefly review it. We shall consider the Lure system given by (6.1)\u2013(6.3), with the linear part satisfying the condition (6.4). Let us assume that the output signal can be described as a transient oscillation that has variable frequency and amplitude. The geometric representation of this motion is a rotating phasor (Fig. 6.1) which is a vector whose length corresponds to the amplitude and the angle to the phase of the output y(t): where a represents the length of the phasor and \u03a8 represents the angle between the phasor and the real axis. We can associate either the real or the imaginary part of the phasor y\u0304(t) with the real signal y(t). Therefore, a(t) = |y\u0304(t)|, \u03a8 (t) = arg y\u0304(t). Let us assume that the initial values of the instantaneous amplitude and the instantaneous phase angle are a(0) and\u03a8 (0), respectively, and introduce the variable \u0393 (t) such that a(t) = a(0)e\u0393 (t)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure10-1.png",
        "caption": "Fig. 10 Push-pulling the axis support model and the modeling results: a an ordinary push\u2013pull edit; and b an arbitrary push\u2013 pull edit",
        "texts": [
            " From the case studies conducted, it has been found that Siemens NX has a better performance in terms of preserving G1 connections; ANSYS SpaceClaim sometimes the former even gives nonsense model shapes. (ANSYS SpaceClaim, however, outperforms Siemens NX for push\u2013pull without G1 connections.) There are in total eight case studies, which are based on six real-world mechanical parts obtained from the GrabCAD part library (https:// grabc ad. com/ libra ry). Case study 1 considered push-pulling an axis support part model where no critical points could be detected during the push\u2013pull edit (Fig.\u00a010). Case study 2 involved a connecting rod part model with one critical point in the push\u2013pull edit (Fig.\u00a011). Case study 3 analyzed push-pulling a hook part model with one critical point in the push\u2013pull edit (Fig.\u00a012). The varied modeling results in these three case studies will be used to show the essential role GTI detection plays in attaining robust push\u2013pull with G1 connections, as well as the effectiveness of the proposed GTI detection method. Case studies 4\u20137 considered three comprehensive modeling examples where various push\u2013pull edits were applied (Figs.\u00a013, 14, 15, 16); they will be used to show the effectiveness of the proposed method as a whole. Case study 8 (Fig.\u00a017) is intended to show some important limitations of the proposed method. In case study 1, there was no critical point in the push\u2013pull edit. Model update was thus made trivial and only involved model regeneration. Siemens NX, SpaceCl and the proposed method successfully gave satisfactory modeling results (Fig.\u00a010), even when the push\u2013pull was made wild (Fig.\u00a010b). In case study 2, similar faces to those in case study 1 were push-pulled, but an invalid modeling result was generated in Siemens NX, and a valid yet unpredictable result (as indicated by the red circle) was given by SpaceClaim (Fig.\u00a011). (Siemens NX colors boundary faces in red whenever there is a model update failure, as shown by the circled face in Fig.\u00a011.) The major difference between case studies 1 and 2 is that the latter involves a critical point of GTI. It can thus be concluded that crossing critical points could cause model update failures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure1-1.png",
        "caption": "Fig. 1: Overview of calculation approach",
        "texts": [
            " Ponick are with Institute for Drive Systems and Power Electronics at Leibniz Universita\u0308t Hannover, Germany (stephan.vip@ial.uni-hannover.de, jan.andresen@ial.uni-hannover.de, ponick@ial.uni-hannover.de). F. Dra\u0308ger is with Robert Bosch GmbH, Germany (florian.draeger@de.bosch.com). the geometry of the electrical machine, the acoustic behaviour needs to be investigated at an early stage of the design process. The process of noise prediction of an electrical machine couples the physical domains electromagnetics, structural dynamics and acoustics [4]\u2013[7]. Fig. 1 presents this cause-and-effectchain. Starting with the electromagnetic calculation, which can be done by FEA, by reluctance networks or in the case of induction machines and salient pole synchronous machines analytically, the spatial force harmonics \u03c3(\u03b3\u2032, t) are derived [4], [8]\u2013[11]. Secondly, the structural dynamics are analyzed by means of a modal analysis. Using modal superposition [12], the surface deflection \u03be is determined for each spatial harmonic acting on the structure. Finally, the emitted acoustic noise can be estimated by an analytical radiator model [13] or by boundary element method (BEM) [14]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001503_embc44109.2020.9176089-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001503_embc44109.2020.9176089-Figure1-1.png",
        "caption": "Fig. 1. (a) The studied tendon-driven limb in MuJoCo environment. (b) each musculotendon consists of a muscle model (M), elastic element (K), and a damper (B).",
        "texts": [
            " Our results also show that an appropriate value of added stiffness can enhance the learning and precision in all cases and even exhibit emergence of lower energy consumption. This is of great significance because the elasticity that is inherent to some types of plants (i.e., tendondriven systems) can now be leveraged to improve learning and performance. In this paper, we studied how adding elastic elements affects autonomous learning in a two-joint three-tendons simulated limb (similar to the physical system introduced in [23], [25]) in the MuJoCo environment [32](Fig. 1.a). The muscle model we use consist of a contractile element with Force-Length-Velocity properties [20], [32], a small parallel damper (100 Ns/m) and a parallel elastic element with stiffness value \u2018K\u2019 (see Fig. 1.b). Specifically, we studied the convergence of the inverse kinematics map, how its performance accuracy changes with stiffness, as well as its adaptability when learning with one stiffness value and then performing using a different value. As for learning, we used our autonomous few-shot hierarchical learning algorithm General-to-Particular (G2P) [23], and the end-to-end Proximal Policy Optimization (PPO) autonomous learning algorithm [33]. G2P is a hierarchical autonomous learning algorithm that, on its lower level, creates an inverse kinematics map using output kinematics collected from an initial random set of actuation commands (motor babbling)",
            " Systems that perform end-to-end learning (such as PPO), on the other hand, usually require larger number of samples to learn to perform a task, are harder to interpret due to their implicit modeling, and usually cannot generalize well across tasks [33], [38]\u2013[41]. These methods, however, can achieve better asymptotic performance even in challenging tasks. For this study, we have performed three set of simulated experiments. In all simulations, elastic elements are considered as parallel elements with each musculotendon (Fig. 1.b); the stiffness value of all elements are equal for each simulation and referred to as \u201cstiffness\u201d. The details for each of these set of simulations are provided below. 1) Controlling the limb with different stiffness values in the muscle model: In this simulation, for each stiffness value, we first randomly activated muscles and recorded the resulting kinematics (motor babbling [23]) for 3 minutes (100 samples per second). The recorded kinematics are joint angles, angular velocities, and angular accelerations for both joints (a vector of 6 values)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001218_ec-10-2019-0447-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001218_ec-10-2019-0447-Figure1-1.png",
        "caption": "Figure 1. Schematic of the DED simulation modeling",
        "texts": [
            "s), which does not represent the viscosity of the solid metal. The material addition in the melt pool is modeled by a boundary condition on the free surface in the material addition direction (z). This condition is represented by a velocity Vz x; y\u00f0 \u00de depending on the effective powder feeding rate Dm such that (Peyre et al., 2017; Arrizubieta et al., 2018): Vz x; y\u00f0 \u00de \u00bc Dm rpR2 1 x2 \u00fe y2 R2 1 2 for x2 \u00fe y2 #R (7) whereR is the powder steam radius. The following thermal boundary conditions are applied on the boundary of the domain (Figure 1): q \u00bc l T\u00f0 \u00de @T @n \u00bc d ij acosu pR2 0 Plaser hc T T1\u00f0 \u00de \u00abs T4 T4 1 (8) where d ij is the Kronecker symbol, which is 1 on the free surface and 0 elsewhere, Plaser is the power of the laser beam, u is the incidence angle to the normal to the substrate n, R0 is the laser beam radius, a is the absorptivity of the material, hc is the heat convection coefficient, EC \u00ab is the emissivity, s is the Stefan-Boltzmann constant (s = 5.67 10 8 W/m2/K4) and T1 is the ambient temperature (T1= 25\u00b0C). The conservation equations (1)\u2013(3) are solved by the FPM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002753_s00202-021-01301-w-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002753_s00202-021-01301-w-Figure8-1.png",
        "caption": "Fig. 8 Magnetic flux density of in-wheel PM BLDC motor",
        "texts": [
            " Among these designs, the 24 slot / 20 pole design, which has the same power value with the same parameters, Table 2 Impact to efficiency, power and torque of in-wheel PM BLDC with the segmented magnets Total number of axial magnet segments used in calculation of magnet losses Efficiency [%] Power of Shaft [Watt] Torque of Shaft [Nm] 24/20 Slot/Pole 36/30 Slot/Pole 42/36 Slot/Pole 24/20 Slot/Pole 36/30 Slot/Pole 42/36 Slot/Pole 24/20 Slot/Pole 36/30 Slot/Pole 42/36 Slot/Pole Fig. 5 Effect of axial magnet segments on efficiency of in wheel PM BLDC motor Fig. 6 Effect of axial magnet segments to power of in wheel PM BLDC motor 1 3 has the highest efficiency. After that two-dimensional Finite Element Analysis carried out for the optimum designed 24 slot / 20 pole motor geometry and magnetic flux density is shown in Fig.\u00a08. The usage of magnetic core is determined by the proper flux distribution within specific limit values. The near saturated sections whose flux density is around 2\u00a0T are shown in the edges of teeth only. The magnetic flux lines distribution is homogeneous and smooth as given in Fig.\u00a09. The certain knowledge of the magnetic flux lines distribution taking these factors into account is essential for the accurate prediction of the motor performance. When the design parameters of the motor are determined properly, this specified magnetic flux line distribution should be homogeneous"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002250_s12666-020-02152-y-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002250_s12666-020-02152-y-Figure8-1.png",
        "caption": "Fig. 8 Tooling (pattern and core) design in CAD environment",
        "texts": [
            " The allowances can be combined and optimized to reduce the total increase in weight and machining required to obtain the desired component. The above routines required a regular CAD program. For this purpose, the team selected SolidWorks, since it was widely used and preferred by foundries owing to its userfriendly features for 3D modeling. Using the underlying routines of the software, a new module called AX-Tooling was developed during 2015\u20132020, to semi-automatically convert a component model into corresponding tooling models (Fig. 8). Some features of the methoding design and casting simulation mentioned earlier were also included in the module. This was meant for product designers in OEMs to quickly assess and improve the design for castability [19]. For example, they could minimize hot-spots by redesigning thick junctions, leading to improved quality and yield, as well as reduced weight and machining cost. This was enabled by algorithms for automatic computation of wall thickness [20] and classification of junctions [21]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003050_01423312211025331-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003050_01423312211025331-Figure2-1.png",
        "caption": "Figure 2. The HiMAT vehicle.",
        "texts": [
            " The observer gains matrix L1s(t) and L2s(t) are chosen such that Gs(t)L1s(t)Bs(t) in (9) and L2s(t)Es(t) in (12) are Hurwitz, respectively. The constant kp is chosen that 1+ kp . 0 must be applied. In the next part, the proposed control method is applied to an example in order to show the performance the method. In this section, we apply the proposed composite adaptive controller consisting of MRAC and DOBC approaches based on proposed switching laws to HiMAT vehicle, which has been studied in (Yuan et al., 2018). Consider the three-view drawing of the HiMAT vehicle as follows. According to Figure 2, the components of the state vector x(t), with the dimension of R3, are as follows: x1(t) is angle of attack, x2(t) is pitch rate and x3(t) is pitch angle, and the components of the control vector u(t), with the dimension of R4, are as follows: u1(t) is elevator, u2(t) is elevon, u3(t) is canard and u4(t) is symmetric aileron. The parameters of HiMAT vehicle based on (1) and (2) are given by: A1 = 0:8435 0:97505 0:0048 8:7072 1:1643 0:0026 0 1 0 2 64 3 75, B1 = 0:1299 7:6833 0 0:092 4:7974 0 0:0107 4:8178 0 0:0827 5:7416 0 2 64 3 75 A2 = 1:8997 0:98312 0:00073 11:720 2:6316 0:00088 0 0 0 2 64 3 75, B2 = 0:2436 46:206 0 0:1708 31:604 0 0:00497 22:396 0 0:1997 31:179 0 2 64 3 75 A3 = 1:2206 0:99411 0:00084 64:071 1:8876 0:00046 0 1 0 2 64 3 75, B3 = 0:0662 27:333 0 0:0315 13:163 0 0:0141 11:058 0 0:0749 26:878 0 2 64 3 75 E1 = 0:1 0:2 0:1 0:1 0:2 0:1 0:2 0:1 0:1 0:2 0:1 0:1 2 64 3 75, E2 = 0:1 0:2 0:1 0:1 0:2 0:1 0:2 0:1 0:1 0:2 0:1 0:1 2 64 3 75, E3 = 0:1 0:2 0:1 0:1 0:2 0:1 0:2 0:1 0:1 0:2 0:1 0:1 2 64 3 75 G1 = 0:01 0:01 0:01 0:01 0:01 0:01 0:01 0:01 2 664 3 775,G2 = 0:01 0:01 0:01 0:01 0:01 0:01 0:01 0:01 2 664 3 775, G3 = 0:01 0:01 0:01 0:1 0:01 0:01 0:01 0:01 2 664 3 775 F1 = 1 1 1 2 ,F2 = 2 1 0 3 ,F3 = 2 1 2 1 H1 = 2 5 ,H2 = 2 5 ,H3 = 2 5 and the switched disturbance observer gains are as follows: L1 1 = 0:29 0:1 0:3 0:1 0:29 0:5 ,L1 2 = 0:29 0:2 0:1 0:1 0:29 0:3 , L1 3 = 0:29 0:5 0:4 0:1 0:29 0:1 L2 1 = 40 10 20 20 90 20 40 20 40 10 40 5 2 664 3 775, L2 2 = 40 10 20 20 90 20 40 20 40 10 40 5 2 664 3 775, L2 3 = 40 10 20 20 90 20 40 20 40 10 40 5 2 664 3 775 where d2 t\u00f0 \u00de= 2 sin (t)e 0:05t; 0:15 sin (pt)e 0:05t;\u00bd sin (pt)e 0:05t; 0:2 sin (5t)e 0:05t and d3 t\u00f0 \u00de= t sin (t)e 0:06t are the external disturbances"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure10-1.png",
        "caption": "Fig. 10. Flux density distribution of proposed RPS-HTSFS machine: (a) Excited only by HTS-excitation field and (b) Excited only by armature field.",
        "texts": [
            " The key design parameters of the proposed machine, compared with the conventional HTS-FS machine, are listed in Table I. The armature winding can produce the magnetic field when the armature current flows in, which can make a disturbance to the HTS-excitation critical field state. Moreover, the over strong armature magnetic field will increase the quench risk Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 11:27:00 UTC from IEEE Xplore. Restrictions apply. of HTS-excitation coil. Fig. 9(a) and Fig. 10(a) show the flux density distributions of the analyzed HTS-FS machines excited by only HTS-excitation field. It can be observed that their flux densities on the U-shaped core are both up to 2.0 T, which can take full use of the strong magnetic field characteristics of the HTS excitation. However, when the armature current flows in, the armature windings can produce the magnetic field, which can make a disturbance to the HTS-excitation critical field state. Moreover, the over strong armature magnetic field will increase the quench risk of HTS coils. Fig. 9(b) shows the flux density distribution of the conventional HTS-FS machine excited by only armature current. Since the armature windings and HTSexcitation coils are placed on the same stator, the magnetic fluxes are all closed by the U-shaped core, causing the HTS coils easily affected by the armature reaction. As a comparison, Fig. 10(b) shows the flux density distribution of the proposed RPS-HTSFS machine excited by only armature current. Thanks to partitioned stator structure, it can realize the physical isolation of armature windings and the HTS-excitation coil. Hence, the variations of magnetic field around the HTS-excitation coil placed on the outer stator are relatively slight, thus improving the operating reliability. Fig. 11 shows the no-load back-EMF waveforms of the proposed RPS-HTSFS machine when the HTS-excitation current is 100 A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure52.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure52.5-1.png",
        "caption": "Fig. 52.5 Modified wheel",
        "texts": [
            " In hilly terrains manoeuvring power tillers or even vehicles for that matter is a challenge due to factors such as steep ascend or descend, poor road conditions, narrow roads, lack of motorable roads and landslides, etc. During rainy seasons, it becomes evenmore dangerous for operators of suchmachines. Therefore, a modified wheel design is proposed for the power tiller so that it can provide sufficient traction while working on such terrains. This wheel design is simple, easy to assemble or disassemble. The CADmodel of the wheel is shown in Fig. 52.5. The circumference of the wheel is provided with lugs of the same width as that of the wheel so that it \u201cbites\u201d into the soil to provide traction without skidding while forming new terrace fields. Agriculture is the basic livelihood of the people in North-Eastern Region (NER) of India. The region consists of both plain and hilly areas which include inaccessible terrain, fragility, marginality, excessive sloping land with rolling topography, rich biodiversity, unique ethnicity and socio-ecological set-up"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure21-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure21-1.png",
        "caption": "Fig. 21. Demagnetization ratio distribution of 10-pole/12-slot STPSPM machine with static eccentricity. (a) Eccentricity ratio = 0.1. (b) Eccentricity ratio = 0.3. (c) Eccentricity ratio = 0.7. (d) Eccentricity ratio = 0.9.",
        "texts": [
            "org/publications_standards/publications/rights/index.html for more information. E. Influence of Eccentricity Ratio The influence of eccentricity ratio on demagnetization and post-demagnetization UMF is investigated in this part, where the symmetrical SPM machines with static eccentricity is investigated as an example. When the eccentricity ratio increases, the PMs facing larger airgap suffer from severer demagnetization whilst the demagnetization of PMs facing smaller air-gap is further mitigated, as shown in Fig. 21. As aforementioned, for the symmetrical SPM machines with static eccentricity, the demagnetization reduces the constant component and the 10th order harmonic and adds the 1st order harmonic. When the eccentricity ratio increases, the constant component and the 10th harmonic decrease more whilst the additional 1st order harmonic increases, as shown in Fig. 22-Fig. 24. Therefore, the variations of UMF harmonics due to demagnetization increase with eccentricity ratio. Authorized licensed use limited to: California State University Fresno"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001863_s12239-020-0147-z-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001863_s12239-020-0147-z-Figure13-1.png",
        "caption": "Figure 13. Schematic of key points on piston set-liner.",
        "texts": [
            "5 \u2103 due to the friction heat, and the effects of friction heat are primarily focused on the area near ring pack as well as the sliding region of piston ring in the inner surface of liner. Accordingly, the largest temperature increase caused by friction heat is some 5 \u2103 on liner face. In terms of the piston, the impact of friction heat decreases gradually with increasing the distance from ring pack because the allocated friction heat of piston ring is directly exerted on ring pack. Figure 12 shows that the maximum temperature wave caused by friction heat occurs on piston ring surface. Therefore, the key points in Figure 13 (angular and central points) are selected to analyze the temperature waves considering the effects of friction heat, as shown in Figure 14. It can be seen from Figure 14 that the effects of friction heat gradually decrease with increasing the distance from the outer surface of piston ring (for instance, points c and b of the top ring), and especially the temperature fields of the inner surface of the top ring do not ultimately change with crank angle, this is because the friction heat is directly exerted on the outer surface of piston ring and gradually permeate from the outside"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002703_j.optlastec.2021.107100-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002703_j.optlastec.2021.107100-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the processing of TWBs. (a) Precutting. (b) Butt welding.",
        "texts": [
            " Based on this novel method, 304SS TWBs with the thickness combination of 2.0 mm and 1.0 mm are successfully butt welded by using DLS. Welded samples are free from cracks and pores. Through tensile strength tests, all welded samples fracture at the thin base metal plate. This work not only demonstrates the processing effect of this novel method, but also provides the experimental guidance for the further research of the diode laser processing. In this work, a novel processing concept of the eave-like model is proposed to realize the welding of 304SS TWBs. Fig. 1 is the schematic diagram of the processing procedure. Firstly, a laser cutting head, tilted at an angle of 80\u25e6 to the work surface, is used to precut 1 mm and 2 mm 304SS TWBs for achieving an inclined edge on the metal plates. Then, 304SS metal plates after precutting are butt jointed together, which can form an eave-like structure. Finally, a perpendicularly placed laser welding head is employed for TWBs welding. The marginal area of the thick plate is melted by the laser beam, and the melted liquid flows into the gap"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000922_pedstc49159.2020.9088409-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000922_pedstc49159.2020.9088409-Figure8-1.png",
        "caption": "Fig 8: The virtual boundary",
        "texts": [
            " The relation between turn function and winding function is as below: ( ) = ( ) \u2212 < ( ) > (1) where ( ) , ( ) , and < ( ) > are winding function, turn function, and mean of turn function, respectively [18]. The winding functions of abovementioned machine are shown in Fig 7. It is assumed that the upward current flow and downward current flow lead to positive and negative Magneto Motive Force (MMF), respectively. Due to the fact that VFRM is a doubly salient machine, the airgap function is not constant when the rotor rotates. In order to achieve more accurate model, a virtual boundary is placed in the machine structure in such a way that Fig.8. shows it. The airgap consists of two functions by this assumption. Fig.9 and Fig.10 show the airgaps which relate to stator and rotor section. According to Fig 9 , the fourier formulation of stator air-gap function is given by : (0) = . + ( \u2212 ). (2) = 2( \u2212 ) . [sin 2\u2212 sin( \u2212 2 )] (3) where , , , , and ( ) are stator static position, the arc of the stator tooth, inverse of the lowest value of the stator airgap function, inverse of the highest value of the stator airgap function, and fourier series of stator air-gap function, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002924_j.msea.2021.141535-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002924_j.msea.2021.141535-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of DLD samples process and the fabricated samples geometry: (a) Schematic diagrams of the DLD samples process, (b) DLD samples with different EAD, (c) The tensile test boards, (d) Geometry of the tensile test boards.",
        "texts": [
            " The 5 fabricated samples with 5 parameters of the experimental group were used to study the evolution mechanism of carbides and to modify the kinetic equation, and the fabricated samples with 4 parameters of the verification group were used to verify the modified precipitation kinetic equation of carbide. EAD was calculated according to equation (1): EAD=P/(D \u22c5 v) (1) where P is the laser power, D is the spot diameter and v is the scanning speed. The overlap rate refers to the percentage of the overlap width between two single track to the width of single track. Q235 was used as the matrix to support the DLD 50Cr6Ni2Y alloy steel samples with a size of 45 mm \u00d7 18 mm \u00d7 6 mm, and the schematic diagram of DLD process is shown in Fig. 1(a). Considering the influence of the matrix on the samples, only the as-deposited part of the initial samples was cut for testing and analysis to study the evolution mechanism and precipitation kinetics of carbides, as shown in Fig. 1(b). The cross-section of the samples were taken for observation and analysis. Fig. 1(c) and (d) display the geometry and size of tensile test boards cut from the samples in Fig. 1(b). The thickness between layers was about 0.5 mm, the laser spot was a rectangular spot of 4 \u00d7 4 mm, and argon protection was provided during printing. The DLD 50Cr6Ni2Y alloy steel samples were ground with sandpaper of different roughness; and then polished with 1.5 m of diamond paste. A corrosive solution (4 mg CuSO4+20 ml HCl+20 ml H2O) was used for 5\u201315 s. The samples were put in an ultrasonic cleaner for 20 min. The electrolytic extraction experiment was used to extract carbides of 50Cr6Ni2Y alloy steel samples, and the electrolyte was a 10% hydrochloric acid and methanol solution",
            " The AG-X100kN electronic universal material testing machine was used to measure the mechanical properties of the deposited samples, and the tensile speed was 5 mm/min at room temperature. In the DLD manufacturing process, directionality is an important issue that impacts X. Chen et al. Materials Science & Engineering A 820 (2021) 141535 the grain structure, texture, size, composition and mechanical properties [22,23]. In the tensile test, the tensile boards were cut parallel to the deposition direction in the fabricated samples, as shown in Fig. 1(c) and (d). Under the same experimental conditions, the tests were carried out three times to ensure the accuracy of the experimental data. Fig. 2 shows the metallographic photos of DLD 50Cr6Ni2Y alloy steel samples under different EAD. It can be seen that the five samples have good shaping, and a uniform metallurgical bond is formed between the layers. The morphology of samples are columnar crystals and equiaxed crystals [24]. Fig. 2(a) shows that when the EAD is 90 J/mm2, there are some pores at the joint between the layers, because the pores along the boundary of the layer were caused by insufficient melting [25]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure3.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure3.6-1.png",
        "caption": "Fig. 3.6 Schematic representation of the single compaction process carried out during the powder metal processing. (a) The metal powder is placed in a die. (b\u2013c) A load is applied to consolidate the loose powder. (d) A green product is obtained",
        "texts": [
            " This process leads to the production of more uniform and edge-free particles in their morphology and whose quality can reach a high degree of spheroidization and batch uniformity. In the case of components made from particles with a high degree of spheroidization, the mechanical bonding among particles, product of the compaction stage, leaves a greater number of interstices among them and, therefore, a greater porosity [13]. 3 Powder Metallurgy 37 Regardless of the route of production of metal powders, the powder metallurgy process involves the mixing and compression of particles in a certain way, as well as the use of lubricants and binders. Figure\u00a03.6 outlines the single cold compaction process. The powders are first loose with no strength and a large number of voids (Fig.\u00a03.6a). As pressure is applied (Figs.\u00a03.6b-c), the particles rearrange, and the contact among them increases. The increment in pressure is responsible for a better packing with the subsequent decrease in porosity, as well as improved intimacy among particles that increases the powder density (Fig.\u00a03.6d). The applied pressure to form the cold compacts should be above the yield strength of the powders. The compaction process also serves to induce elastic and in some cases plastic deformation, which induce strain hardening of the compact and promote recrystallization that is critical for sintering. The consolidation step is conducted to allow proper handling of the samples prior to the sintering process. As an example of the cold compactions process, Fig.\u00a03.7 shows the steps followed for the compaction of aluminum powders reinforced with carbon nanotubes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001147_j.jmmm.2020.167119-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001147_j.jmmm.2020.167119-Figure3-1.png",
        "caption": "Fig. 3. Equivalent circuit of polyphase machines when m > n.",
        "texts": [
            " (38) can be rewritten as follows: = + + + = + + +( ) V I r jX jX I I I jX jX I I ( ) ( ) 0 ' ' ( ) s s s ls m m s r r r s lr m m s r 2 ' ' 2 'r (39) Case 2, in this case, the No. of phases in the stator is more than the No. of phases in the rotor. The stator and rotor voltage equation can be rewritten as in Eq. (40). Voltage \u201cVSI\u201d is an induced voltage from the stator due to the increased number of stator phases; =V jX ISI n m m r2 ' , also an additional inductance is presented in the stator circuit equal to Xm n m2 . The equivalent circuit can be represented as in Fig. 3. Case 3, the No. of rotor phases is more than the No. stator phases. The equivalent circuit voltage equation can be rewritten as in Eq. (41). Additional voltage \u201cVRI\u201d is induced from the rotor side in the stator side; =V jX IRI m n m s2 , also an additional inductance is presented in the rotor side equal to Xn m m2 . The equivalent circuit can be represented as in Fig. 4. Fig. 5. Simulation model operation flowchart. Fig. 4. Equivalent circuit of polyphase machines when m < n. = + + + + = + + + +( ) V I r jX jX I I jX I I jX jX I I jX I ( ) ( ) 0 ' ' ( ) s s s ls m m s r n m m r r r s lr m m s r n m m r 2 ' 2 ' ' 2 ' 2 'r (41) To validate the analytical calculation in this paper, polyphase IM simulation is carried using two software \u201cFinite Element Method Magnetics (FEMM) and MATLAB\u201d"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure1-1.png",
        "caption": "Fig. 1. (a) Major components of orbit motor and configuration at orbit angles of (b) 0 , (c) 90 .",
        "texts": [
            " Acceleration m=s2 b Face width of the gears m\u00f0 \u00de cr Coefficient of Restitution \u00f0 \u00de D Damping Coefficient E Young\u2019s Modulus Pa\u00f0 \u00de e eccentricity m\u00f0 \u00de F Force N\u00f0 \u00de ht Gap Height m\u00f0 \u00de K Contact Stiffness Parameter L Projection area in plane, S b m\u00f0 \u00de _m Mass Flow Rate kg=s\u00f0 ) N Number of Outer Gear Teeth \u00f0 \u00de n Nonlinear Power Exponent \u00f0 \u00de p Pressure Pa\u00f0 \u00de pp Pressure of the inlet/outlet environment Pa\u00f0 \u00de Q Flow rate m3=s R Radius of Curvature m\u00f0 \u00de r Radius m\u00f0 \u00de S Projection of Areas m2 T Torque N m\u00f0 \u00de t time s\u00f0 \u00de V Volume m3 Vd Displacement Volume m3=rev x Arithmetic Mean \u00f0 \u00de Y Yield Strength of Teflon Pa\u00f0 \u00de Greek letters a Orifice Coefficient (\u2013) x Angular Velocity rpm\u00f0 \u00de l Dynamic Viscosity Pa s\u00f0 \u00de m Poisson\u2019s Ratio \u00f0 \u00de r Standard Deviation \u00f0 \u00de d Relative Indentation m\u00f0 \u00de _d Relative Normal Contact Velocity m=s\u00f0 \u00de _d \u00f0 \u00de Velocity of Approach m=s\u00f0 \u00de v Hysteresis Damping Factor \u00f0 \u00de g Efficiency (%) q Density kg=m3 b Ratio of Average Contact Pressure to Yield Strength \u00f0 \u00de k1; k2 Friction Coefficients \u00f0 \u00de Subscripts b Bottom Surface c Normal Contact f Friction ger Gerotor max Maximum meas Measured mech Mechanical N Normal orb Orbit Motor sim Simulated t Top Surface vol Volumetric Abbreviations COV Coefficient of Variation CP Average Contact Pressure EM Electric motor attached to the orbit motor HP Hydraulic Pump PRV Pressure Relief Valve PM Electric Motor attached to the Pump TM Transmission Box TSV Tooth Space Volume Others F Filter Ps Pressure Transducer Qf Flow Meter Tc Thermocouple Ts Torque/Speed Sensor figuration at two different angles are presented in Fig. 1. The cylindrical rollers are represented as red circles enveloping the rotor (blue profile). Gerotors and orbit motors differ mainly in the displacement volume for a given size of machine. The rotor-stator set of orbit motor acts as a gear reducer with gear ratio 1:N, where N is the number of stator teeth (rollers). The displacement volumes of gerotor and orbit motor can be represented as [1]: Vd;orb \u00bc N Vd;ger \u00f01\u00de Due to the high displacement realized by orbit motors, they are particularly convenient for high specific torque applica- tions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002509_14644193211003777-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002509_14644193211003777-Figure3-1.png",
        "caption": "Figure 3. Location of the multi-disk: (a) one disk; (b) three disks; (c) five disks; (d) seven disks.",
        "texts": [
            " These energy terms are brought into the Lagrange equation of the corresponding coordinates to form a set of unified motion equations. This set of unified equations of motion can be represented as equation (4) by the matrix form.27 M\u00bd \u20acqf g \u00fe C\u00bd _qf g \u00fe K\u00bd qf g ~Q \u00bc 0f g (4) where M, C, and K are the mass matrix, stiffness matrix, and damping matrix, respectively. Differential equations of motion for rotor In order to discuss the effect of rotor mass distribution method on the bearing vibration response, the rotor mass is distributed in multi-disk as shown in Figure 3. The rotor mass distribution method in Figure 3(a) is the traditional centralized mass distribution method, in which the mass of the rotor is con- centrated on a thin elastic disk in the center of the shaft. This rotor mass distribution method has the advantage of simple modeling, but it cannot reflect the influence of the actual rotor mass distribution on the system vibration response. Therefore, in Figure 3(b) to (d), the mass of the rotor has been evenly distributed in different number of disks. Both the shaft and disk have two DOFs in X direc- tion and Y direction. The gyroscopic inertial effects have been considered in the disk. A portion of the rotor mass is distributed at both ends of the shaft.30 The kinetic energy of the rotor Tr can be given by equation (5). Tr \u00bc Tti \u00fe Tri \u00bc 1 2 ms\u00f0 _xds \u00fe _yds\u00de2 \u00fe Xn i\u00bc1 mdiski\u00f0 _xdiski \u00fe _ydiski\u00de2\u00fe Xn i\u00bc1 Idiski;t\u00f0 _hdiski;X \u00fe _hdiski;Y\u00de2 \u00fem\u00f0 _xfs \u00fe _yfs\u00de2 (5) where Tti is plane kinetic energy, Tri is rotational kinetic energy, Idiski,t is the transverse moment of inertia of the i-th disk, ms is half of the total shaft mass, and mdiski is the mass of the i-th disk"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001452_icisc47916.2020.9171086-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001452_icisc47916.2020.9171086-Figure1-1.png",
        "caption": "Fig. 1: Frames of references [22]",
        "texts": [
            " For this purpose, PSO is used for its intelligent tuning. The integral part of the controller design is aimed at minimization of tracking error of the UH. The proceeding sections of this paper are organised as: Section II deals with modelling of the UH. Section III states the control objective. Section IV presents the controller design and in section V the simulation results are presented and discussed. To model the helicopter, it is imperative to work in two co-ordinate frames of references, inertial ( NED ) and bodycarried frame. Fig. 1 represents the rigid body diagram of the UH and the co-ordinate frames of reference. From the Newton Euler equations, we arrive at the complete non-linear dynamics of the helicopter, with the exception of 978-1-7281-2813-9/20/$31.00 \u00a92020 IEEE 324 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on August 24,2020 at 04:05:08 UTC from IEEE Xplore. Restrictions apply. equations (10) and (11), which are acquired from literature [21], [22]. u, v, w are the translational velocity components of the UH in body-frame; p, q, r are the rotational velocity components of the UH in body frame; \u03b8, \u03c8, \u03c6 are the Euler angles and a, b are the longitudinal and lateral flapping angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002956_978-3-030-73882-2_156-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002956_978-3-030-73882-2_156-Figure15-1.png",
        "caption": "Fig. 15 The schematic of an H-bridge in EasyEDA software",
        "texts": [
            " Figure 13 shows the schematic of our proposed Fire Safety System in EasyEDA software. And Fig. 14 shows the schematic of the power Supplies system, Voltage Sensor and input pins in same software. We have used the LM317T regulator in schematic to adapt the input power supply voltage that equal 24 V to an adapted voltage at 9 Vwhich used to power the Arduino Uno Card. And also the voltage sensor used to measure the voltage at the resistor terminal. In the power part, the siren and the polarity resistor located in the middle of the H-bridge as illustrated in Fig. 15. In the control part, the Arduino Uno Card with the connection connectors and all the components as well as the output pin and also the part who declares to the proposed Fire Safety System if there is a fire and the status of the siren, are illustrated in Fig. 16. Design and Realization of Fire Safety System \u2026 441 Fig. 16 The schematic of the control part in EasyEDA software We have used a 24 V DC power supply for the siren which powered during a fire. When there is no fire the voltage sensor will measure a zero voltage which means a disconnection for the proposed Fare Safety system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002570_iros.2011.6048138-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002570_iros.2011.6048138-Figure6-1.png",
        "caption": "Fig. 6. Deformation measurement points of the surface of the foil type stator",
        "texts": [
            "000014 \u00b7 sin(2\u03c0 \u00b7 26800 \u00b7 t) (1) This equation is the same as for the wave generator using in a real experimental setting. Fig.5 shows the result of the transient response analysis. First, a longitudinal wave is propagated to the foil stator and then the stator deforms in the radial direction. Deformation of the stator is transmitted from the joint of the stator and waveguide to the tip of the stator. To quantify this deformation, the displacement of the stator surface was observed from the side of the waveguide side. The observation points are shown in Fig.6. Figs. 7, 8, 9 show the analysis result of surface displacement in the y-z coordinate for each turn of the foil. Except for the bottom side, the result shows every observation point rotating in the winding direction of the foil. Lamb wave usually consists of many modes. but in low ultrasound frequency like our experimental situation, it has mainly two modes, symmetric (S0) mode and asymmetric (A0) mode [8]. The surface of the symmetric mode rotates forward direction, and the symmetric mode rotates reverse direction (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure10-1.png",
        "caption": "Fig. 10. Rig under tilting load.",
        "texts": [
            " Similarly, when the outer ring displaced z\u03b4 in z direction, the normal contact deformation ni\u03b4 is expressed as ' 0 0ni A A\u03b4 = \u2212 ( )2 2 1 2 0 sinz id d A\u03b4 \u03c6= \u2212 + \u2212 . (16) It is defined here that when the outer raceway moves to the bearing center, the displacement is positive (squeezing the steel ball), and when the outer raceway moves away from the bearing center, the displacement is negative (out of contact). Therefore, the terms cosy i\u03b4 \u03c6 and sinz i\u03b4 \u03c6 have an opposite sign in Eqs. (15) and (16). When the ring is displaced under tiling load, assume that the tilt angle of the ring is z\u03b8 , as shown in Fig. 10. The outer raceway displacement zxi\u03b8\u03b4 and zyi\u03b8\u03b4 can be expressed as ( ) cos 2 2 1 cos 2 sin 2 mi m i mi zyi z mi zxi z d d d d \u03b8 \u03b8 \u03c6 \u03b4 \u03b8 \u03b4 \u03b8 \u23a7 = \u00d7\u23aa \u23aa \u23aa = \u2212\u23a8 \u23aa \u23aa =\u23aa \u23a9 . (17) The relationship of the raceway curvature center is shown in Fig. 11. So, the normal contact deformation ni\u03b4 between the steel ball and the raceway is, ' 0 0ni A A\u03b4 = \u2212 '2 ' 2 1 2 0d d A= + \u2212 . (18) ( ) ( ) 2 2 1 2 0zyi zxid d A\u03b8 \u03b8\u03b4 \u03b4= + + + \u2212 The relationship between the tilt displacement. z\u03b8 of the inner ring and the normal contact deformation ni\u03b4 can be established by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000184_epe.2019.8914882-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000184_epe.2019.8914882-Figure5-1.png",
        "caption": "Fig. 5: Demagnetization ratio distribution of the proposed PMASynRM (Ie = 180A, \u03b2 = 90\u00b0). (a) 100\u00b0C. (b) 150\u00b0C.",
        "texts": [
            "1 ( )2 1 1 100 %r r B B \u03b4   = \u2212 \u00d7    (1) In addition, the performance after irreversible demagnetization is investigated on the basis of demagnetization analysis, by considering each element of the PMs as a PM having a new demagnetization curve (Br = Br2, slope \u03bcrec) in the 2D finite-element software. This study assessed the worst demagnetizing conditions (the phase current Ie of 180 A and the current phase angle \u03b2 of 90\u00b0), which do not occur during normal operation [6]. Ie is the rms value, and \u03b2 represents the leading angle of the current vector from the q-axis. Fig. 5 shows the demagnetization ratio distribution of the bonded magnets in the proposed PMASynRM at 100\u00b0C and 150\u00b0C. In the case of 150\u00b0C, the demagnetizing area was expanded compared to the case of 100\u00b0C because the coercivity of the magnets decreased owing to the temperature-rise as shown in Fig. 3. Moreover, the air gap side of both the first and second layers, which is more affected by the reverse magnetic field, was most demagnetized. EPE'19 ECCE Europe ISBN: 978-9-0758-1531-3 - IEEE catalog number: CFP19850-ART P"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure9-1.png",
        "caption": "Fig. 9. Top view schematic of 3-R\u0304RR manipulator at home position. It shows the manipulator parameters, such as link lengths, top platform size and joint angle notations. The proximal link is denoted as lp and the distal link as ld. denotes the angle of proximal link of leg i.",
        "texts": [
            " Then, the resulting matrix in (20) represents the transformation of input torques to output forces. \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 fx fy fz m\u03d5x m\u03d5y m\u03d5z \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 lxd1 D1 lxd2 D2 .... lxdn Dn lyd1 D1 lyd2 D2 .... lydn Dn lzd1 D1 lzd2 D2 .... lzdn Dn e1ylzd1 D1 e2ylzd2 D2 .... enylzdn Dn \u2212e1xlzd1 D1 \u2212e2xlzd2 D2 .... \u2212enxlzdn Dn e1xlyd1\u2212e1ylxd1 D1 e2xlyd2\u2212e2ylxd2 D2 .... enxlydn\u2212enylxdn Dn \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c41 \u03c42 . . \u03c4n \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (20) A planar symmetric 3-DOF parallel manipulator (3-R\u0304RR) is shown in Fig. 9. Rbase and Rtop define the size of the fixed platform and mobile platform, respectively; lp1 and ld1 are the lengths of the first and second links of the three legs, respectively. The link lengths are indeed the projected link lengths and they are divided by cos\u03b4 to get the isometric link lengths. Due to symmetric condition, for i = 1, 2, 3; lpi = lp; ldi = ld. The output forces obtained in (20) provide the output transmission capability. Indeed, the parallel manipulators also possess input transmission capability. Hence, an input transmission index (ITI) is determined along with the output transmission index (OTI). For an individual R\u0304RR leg consisting of three joints, the transmission wrench is a pure torque. Thus, the ITI is obtained as the absolute of the sine of angle between the two Z planes,24 each plane containing the input link and the coupler link, respectively. Referring to Fig. 9, the ITI = |sin (\u03b82)i| (21) and the OTI = |sin (\u03b83)i| (22) A local transmission index (LTI) captures both the indices, by finding the minimum of the two indices. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S026357472000020X Downloaded from https://www.cambridge.org/core. San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at The workspace of the individual R\u0304RR leg i is represented by reachable circle whose radius is lp1 + ld1. The center of each individual reachable circle is initially at the corresponding base platform coordinates A1, A2, A3 (Fig. 9). For finding the mutual workspace of all the three legs, the individual workspace of each leg is translated from their base platform to a vector vi of magnitude equivalent to the radius of the mobile platform Rtop, along the direction of mobile platform orientation \u03d5\u25e6 (Fig. 10). The void circle and the reachable circle (of leg 2) are shown as red dotted circles, before the translation and black circles, after the translation by a magnitude Rtop. The reachable circles C1,C2 and C3 corresponding to each leg are translated, and the obtained mutual workspace is denoted as M1,M2 and M3 in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure1-1.png",
        "caption": "Figure 1. Process of beveloid gear surface.",
        "texts": [
            " Obtaining the precise integral range in potential energy calculation formula, which means obtaining the precise location of the integral start point and end point, is the key problem to accurately calculate the potential energy of the gear mesh process. The integral start point can be solved by the analytical expression of beveloid gear. Getting the integral end point will rely on getting the exact contact points, so the theoretical tooth contact model is proposed first. The beveloid gear is processed with gear hobbing technique by an inline hobbing cutter which is shown in Figure 1. The tooth surface equation taking gear 1 as example in Oo xoyozo based on former research10 can be expressed as XO g1 = xccut cosu1 + yccut sinu1 + r1 sinu1 r1u1 cosu1 xccut sinu1 + yccut cosu1 + r1 cosu1 + r1u1 sinu1 zccut z11 2 4 3 5 \u00f01\u00de where xccuty c cutz c cut T are the coordinates of cutter\u2019s sur- face in Oc xcyczc, r1 denotes the reference circle radius for gear 1, z11 denotes profile shift coefficient between processed gear end face and plane Oo xoyo, u1 can be solved by mesh equation which is u1 = xccutv1n c y v1y c cutn c x r1v1ncy \u00f02\u00de where ncxn c yn c z h iT is the normal vector of cutter\u2019s surface, v1 is the assumed gear rough rotation speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure4-1.png",
        "caption": "FIGURE 4. Static deflection contour",
        "texts": [
            " BC must represent the real condition of object. In this paper BC was lied on the bottom part of frame since this part was connected to the chassis of electric bus. The type of BC is fixed which the rotation and displacement are not allowed in all directions. The location of BC was marked by the red color on Figure 2. While the load with the magnitude and angle based on the ECE 66 standard was shown in Figure 3. 030153-3 The output of simulation is static deflection distribution along the frame that shown in Figure 4. From the results of static simulation, it can be seen the value of static deflection on the structure of the electric bus frame. The maximum deflection value is 756,69 mm while for the minimum deflection it is 84.06 mm and the average static deflection value is 420.03 mm. After the static deflection value was obtained from the static loading simulation results, then the impact factor values are calculated. This impact factor will be used as a first step in dynamic simulation. Another output of static simulation is the von Misses stress"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure12-1.png",
        "caption": "Fig. 12. Topology of the \u201cSPM_DC1_DC2_NPM_DC1_DC2\u201d HEBFM.",
        "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. By adding another ferrite tooth between the PM teeth, the operation statuses can exponentially increase. For instance, one DC winding has 3 operation statuses. Then two DC winding will have 32=9 operation statuses. This topology can be nominated as \u201cSPM_DC1_DC2_NPM_DC1_DC2\u201d HEBFM. Configuration of this type of HEBFM is shown in Fig. 12. The configurations of DC windings and PMs in the nine operation statuses are shown in Fig. 13. Since the DC windings around the N pole and S pole are arranged symmetrically, only the N pole part is demonstrated in order to show the flux lines more clearly. In addition, the corresponding flux linkage vectors are shown in Fig. 14. Here, IDC1 and IDC2 represent the current of DC1 winding and DC2 winding, respectively. Based on the derivations from \u201cSPM_DC_NPM_DC\u201d HEBFM to \u201cSPM_DC1_DC2_NPM_DC1_DC2\u201d HEBFM, it can be deduced that if the number of DC coils between two adjacent S and N PM poles is x, there will be 3x different operation statuses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure7.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure7.6-1.png",
        "caption": "Fig. 7.6 Schematic operation principle of the cold spray process",
        "texts": [
            " Cold-sprayed Al coatings can be applied to protect metal surfaces from atmospheric degradation, because a very thin and impervious oxide layer is formed on the aluminum surface. Instead of replacing the whole structure, repair and reclamation of the coating are possible, with the required protection coming from the sprayed aluminum structure. Hence, the cold spray process is highly attractive, enabling tremendous savings by becoming an inherent part of an integral manufacturing process for Al and its alloys. As shown in Fig.\u00a07.6, the main elements of the cold spray setup are the spraying unit consisting of a prechamber and a supersonic nozzle, the powder feeder, the gas heater, and the source of compressed gas. Particles are accelerated to very high velocities by the carrier gas forced through a nozzle. Upon impact, solid particles with sufficient kinetic energy deform plastically and bond mechanically to the substrate to form a coating. The characteristic features of the cold spray process are much lower temperature (600\u00a0\u00b0C) and higher velocity of particles (>700\u00a0m/s) [22]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001215_1748006x20933506-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001215_1748006x20933506-Figure3-1.png",
        "caption": "Figure 3. Location relationship between curvature centers of inner and outer races.",
        "texts": [
            "33 Then, the component force FL can be obtained as FL \u00bc FP nKH\u00f0dL\u00de 3 2 sin aL \u00f0FL \\FP\u00de Fp \u00f0FL \u00f8FP\u00de \u00f03\u00de Similarly, the normal force QR and deformation dR of each right ball as a result of the horizontal force Fa are expressed as equation (4), where aR is the right contact angle QR \u00bc FP +FR n sin aR dR \u00bc 1 KH 2 3 Q 2=3 R 8< : \u00f04\u00de Then, the component force FR can be obtained as FR \u00bc nKH\u00f0dR\u00de 3 2 sin aR FP \u00f05\u00de Considering the horizontal force Fa, the mechanics equilibrium equation of matched bearings is derived as Fa \u00bc nKH dR\u00f0 \u00de 3 2 sin aR nKH dL\u00f0 \u00de 3 2 sin aL \u00f0FL \\FP\u00de nKH dR\u00f0 \u00de 3 2 sin aR \u00f0FL \u00f8FP\u00de ( \u00f06\u00de Figure 3 shows location relationship between curvature centers of inner and outer races. In view of Figures 1 and 3, OBL and OBR are the left and right ball centers, respectively; rI and rO are the curvature radiuses of inner and outer races, respectively; aO is the initial contact angle; OIL, OIR and OOL, OOR are the left and right curvature centers, respectively, of inner and outer races. When matched bearings operating only with the preload FP, curvature centers OIL and OIR will transform to positions O0IL and O0IR, respectively; aP is the new contact angle and dP is the normal deformation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001863_s12239-020-0147-z-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001863_s12239-020-0147-z-Figure2-1.png",
        "caption": "Figure 2. Structure of ring.",
        "texts": [
            " In this case, all elements in the region on piston set have the friction heat relation with the element el i. Finally, the friction heat qlip of the element el i in the time interval tlip is expressed in integral mean form (9) From to , the friction heat qlio of the element el i is zero because of contacting with piston skirt and lubrication oil. The corresponding acting time tlio is evaluated by the contacting positions and . Finally, the friction heat of the element el i from to is (10) As illustrated in Figure 2 (The origin is at the center axis of liner), assume that the offset swing of piston ring does not occur, the angle between the piston ring open and antithrust side is  corresponding to the starting point of the coordinate . Also suppose that the inner radius of liner is R, the position in the circumferential direction of piston ring is expressed as arc length y, and then . 3.1. Average Reynolds Equation Consider the piston ring and liner with cross-section as shown in Figure 3 (for simplicity, the case that piston rings are fully lubricated by lubrication oil only considered in this paper), the average Reynolds equation (Patir and Cheng, 1978) used to solve the pressure distribution in lubrication film is written as (11) where p is the lubricant pressure, x and y are the pressure flow factors, s is the shear flow factor,  is the dynamic viscosity, u is the piston velocity, is the mean film thickness, and is the composite roughness and stated as t t  l a l a l a ( , , )A r x l d l d l d ( , , )D r x l a l a l a ( , , )A r x l b l b l b( , , )B r x l a l a l a ( , , )A r x l b l b l b( , , )B r x l b l b l b( , , )B r x l c l c l c ( , , )C r x l b l b l b( , , )B r x l c l c l c ( , , )C r x p j p j p j( , , )r x l j l j l j l j( , , , )Q r x t a b c d   l ip t l i p l j l j l j l j 0 li p 1 ( , , , )q Q r x t dt t      l c l c l c ( , , )C r x l d l d l d ( , , )D r x l c l c l c ( , , )C r x l d l d l d ( , , )D r x l i q l a l a l a ( , , )A r x l d l d l d ( , , )D r x l i l i g lig li p li p lio lio 1 ( )q q t q t q t t        y R 3 3 x y T Ts 12 12 2 2 h p h p x x y y h hu u x x t                                 T h where 1 and 2 are the roughnesses of ring and liner faces, separately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000167_iccnea.2019.00088-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000167_iccnea.2019.00088-Figure1-1.png",
        "caption": "Figure 1. Stress diagram of X-type quadrotor aircraft model.",
        "texts": [
            " DYNAMIC MODELING OF QUADROTOR AIRCRAFT Firstly, to establish dynamic model of quadrotor aircraft, the assumptions are as follows: 1) In the inertial coordinate system, the acceleration of gravity is constant and ignoring the influence of the curvature of the earth and the earth's rotation. 2) In the process of flight, regarding the quadrotor as a rigid body, and the vibration and deformation are neglected. 3) The center of gravity of a quadrotor coincides with the center of the body structure. 4) The mass and torque of inertia of a quadrotor are constant. 5) Ignoring ground effect and air drag, quadrotor is only pulled by gravity and propeller. The common structures of quadrotor aircraft are X-type and cross-type [8]. This paper chooses X-type quadrotor as the research object. Fig.1 demonstrates the stress diagram of X-type quadrotor aircraft model. 447 978-1-7281-3977-7/19/$31.00 \u00a92019 IEEE DOI 10.1109/ICCNEA.2019.00088 B is the body axis system, and E is the earth fixed axis system. T is defined as a column vector consisting of roll angle, 33RR pitch angle and yaw angle of the body. The matrix is the rotation which is used to relate a vector in the body axis system to earth fixed axis system defined as 11 12 13 21 22 23 31 32 33 R( ) = C C C c c s s c c s c s c s s C C C c s s s s c c c s s s c C C C s s c c c (1) Where the abbreviations c and s have been used for cos and sin respectively",
            " According to NewtonEuler theorem, establish the dynamic equation of quadrotor under the B system as 3 3 3 3 3 3 0 0 b b b b b b b b mI v mv F J J (2) Among which x= T b y zF F F F is the resultant force on the body, x T b y z is the resultant torque of body, T bv u v w defined as linear velocity of aircraft, T b p q r defined as angular velocity of aircraft, m represents total body mass, the three-rank matrix 3 3 1 1 1I diag and 3 3 xx yy zzJ diag J J J represent identity matrix and inertia matrix of body. Based on the model shown in Fig.1 and the above assumptions, define 1 2 3 4 TU U U U U as a matrix of force and torque on body( 1U is the resultant pull of four propellers, 2U , 3U and 4U represents the pitching, rolling and yawing moment, respectively.). According to the blade element theory, the pull iT and reverse torque iQ of the i th propeller are written as 2 i T iT k (3) and 2.i Q iQ k (4) Quadrotor aircraft is subject to gravity G and the pull of four propellers iT ( =1,2,3,4)i during flight. Gravity is expressed as T e mgG 00 under E-system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure4-1.png",
        "caption": "Figure 4. Simulation model of the machine tool in the VERICUT. (a) virtual machine tool CZXB6140 and (b) components tree of the machine tool.",
        "texts": [
            " In particular, the machine tool simulation module can help users complete the real 3D simulation of the entire CNC machine, just like the actual production of the workshop.19 The simulation of the CZXB method is performed in VERICUT.20 According to the process system of the CZXB method in Section \u2018\u2018Process system for machining hourglass worms by the CZXB method\u2019\u2019, the virtual machine includes four CNC axes. Refer to the layout of the horizontal lathe, components tree of the machine tool is built in VERICUT, as shown in Figure 4(b). The tool branch is \u2018\u2018Base>Z>X> B>Tool holder/Tool\u2019\u2019. The stock branch is \u2018\u2018Base>C>Fixture> Stock\u2019\u2019. With references to the structure and dimensions of the C6140 lathe, the 3D model of the main components of the virtual machine tool CZXB6140 are built in the CAD software. These 3D models are exported as STL files and loaded into the corresponding components of the component tree in VERICUT. The initial position of the machine tool and the stroke limit of the CNC axis are set. The virtual machine tool CZXB6140 is shown in Figure 4(a). Call the control system file and select Fanuc oi as the control system of CZXB6140. The simulation model of the machine tool that has CZXB functions is shown in Figure 4. Processing simulation of the hourglass worm According to the process system of the CZXB method in Section \u2018\u2018Process system for machining hourglass worms by the CZXB method\u2019\u2019, tool profiles are not specified in the CZXB method. To make it easier to create a 3D model of the tool and adjust its pose in VERICUT, a tool with a finger tapered profile is selected in this section. Therefore, we will get a finger tapered surface enveloping hourglass worm21\u201326 in the simulation. A finger tapered surface enveloping hourglass worm with multi-threads is described as an example of the CZXB method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001845_s00202-020-01132-1-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001845_s00202-020-01132-1-Figure10-1.png",
        "caption": "Fig. 10 Replacement of the rotor of SynRM with the shaft",
        "texts": [
            " Besides, it considers the effect of the convection coefficient on the amount of temperature of the winding and surface one. Figure 9 shows the diagram of the test setup. Accordingly, to rotate the cooling fan, the shaft of SynRM was coupled to the shaft of an induction motor. For producing different fan rotational speeds, the speed of the inductionmotor is adjusted by a frequency converter. To better identify the convection phenomenon, it is essential to define the system only based on the stator joule losses; accordingly, as shown in Fig. 10, the rotor of the SynRM machine was replaced by a shaft. Then the cooling fan was assembled at the end part of the shaft. Figure 11 shows the test setup. Accordingly, the DC stator experiment is a standard test method to determine the convection coefficient. For this purpose, the three-phase stator windings of the SynRM are connected in series and are supplied through a DC power supply. During the experiments, in addition to the injected power to the stator windings, the temperature of the machine housing, end winding, slots, and inlet flow are measured at different locations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.6-1.png",
        "caption": "Fig. 57.6 Loading in FEA model",
        "texts": [
            " In order to arrive at the boundary conditions, a seat load impression test was carried out using blue print method on existing seat base under the load conditions of: nopassenger, solo passenger, and twin passenger configurations. The boundary conditions were applied at the initial step in the CAD model (Fig. 57.5). The seat base bottom rubber cushion mounting and the rear rubber grommet mounting region were constrained along the vertical direction. A uniform pressure distribution of 18.5 kPa (corresponding to 2600 N of distributed load) was applied over the rider and the pillion seating area in the vertical direction (Fig. 57.6). The failure criteria used for homogeneous materials are not sufficient for predicting failure in composite lamina. This is because the planes along which the lamina may be possibly be the weakest need not be the direction of principal stresses in a lamina. Thus, alternate failure theories have been developed to predict the failure of composite lamina. A non-interactive failure criteria were used as the interactions between the stresses/strains in the lamina is not necessary. The failure modes are predicted by comparing the individual stresses/strains with respect to their ultimate stresses/strains"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure29-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure29-1.png",
        "caption": "Fig. 29. Tilting load-displacement test bench.",
        "texts": [
            " The bearing axial load-displacement test bench is shown in Fig. 27. The bearing inner ring and the shaft are rigidly fixed, and the outer ring is rigidly connected to the housing, as shown in Fig. 28. Axial load is applied to the shaft in both directions, and the axial displacement of the shaft is measured by a displacement sensor. The measurement software records the data from the load cell and the displacement sensor at the same time to obtain the axial load-displacement curve. The bearing tilt load-displacement test bench is shown in Fig. 29. The schematic figure of measurement is shown in Fig. 30. The measured load is the radial load with a certain force arm. The measuring object of the displacement sensor is a measuring rod, and the cylindrical surface of the measuring rod is tan- gent to the shaft and is pulled by a rubber rope to ensure that the rod and the shaft are always in contact. The bottom of the measuring rod is connected with the guide rail to ensure that the measuring rod can move freely in the horizontal direction. The load arm length of the test bench is 200 mm, and the measuring arm length is 150 mm",
            "4 \u03bcm, it can be considered that the dis- placement under the load at this time all comes from the deformation of the fixture. The measured bearing axial loaddisplacement curve will be corrected in \u00b11000 N by deducting the curve in Fig. 33. Note that the load and unload curve in Fig. 33 does not overlap. The reason is that the internal friction of the device itself causes the displacement to lag when returning. When the solid block is subjected to a loading force of \u00b1100 N at the position of the loading arm as shown in Fig. 29, the deformation is shown in Fig. 34 using FEM analysis. According to the analysis results in Fig. 34, the maximum radial displacement of the inner bore surface of the bearing is 0.24 \u03bcm and -0.21 \u03bcm, so a radial displacement of 5.6 \u03bcm will occur at the point of the measuring arm. See Fig. 35. The measured solid block tilt load-displacement curve is shown in Fig. 36. The displacement of the system under the tilt load is 75.7 \u03bcm in the positive direction and -69.2 \u03bcm in the negative direction. The actual measured tilt load-displacement curve will be corrected on the \u00b1100 N range using the curve in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002120_s00170-020-06571-5-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002120_s00170-020-06571-5-Figure10-1.png",
        "caption": "Fig. 10 Main direction of induced normal curvature between cutter and pinion",
        "texts": [
            " Then, through a series of coordinate transformations from oc \u2212 xcyczc to oP \u2212 xPyPzP, the position vector rP and the unit normal vector nP of the point on the pinion flank are obtained, shown as the following equations, rP \u00bc MPPdMPdMMMcr c \u00f045\u00de nP \u00bc mPPdmPdMmMcn c \u00f046\u00de where MPPd and mPPd are the coordinate transformation matrices from oPd \u2212 xPdyPdzPd to oP \u2212 xPyPzP,MPdM andmPdM are the coordinate transformation matrices from oM \u2212 xMyMzM to oPd \u2212 xPdyPdzPd, MMc and mMc are the coordinate transformation matrices from oc \u2212 xcyczc to oM \u2212 xMyMzM, and mPPd, mPdM, and mMc are 3 \u00d7 3 submatrices extracted from the upper left hand corner of MPPd, MPdM, and MMW, respectively. Thus, the first-order parameters of any point on the generated pinion flank are determined. There is a line contact between the cutter and the pinion when the cutter generates the pinion flank. At the meshing point, the first main direction ecP1 of the induced normal curvature between the cutter and the pinion is the tangential direction of the instantaneous contact line, and the second main direction ecP2 is the direction of p, shown in Fig. 10. The principal induced normal curvature KcP 1 along ecP1 , and KcP 2 along ecP2 at the meshing point are as follows: KcP 1 \u00bc 0 KcP 2 \u00bc p\u00f0 \u00de2 ncM \u22c5q\u00fe vcPM \u22c5p 8< : \u00f047\u00de where p \u00bc \u03c9cP M ncM \u00fe Kc 1 vcPM \u22c5ec1M \u22c5ec1M \u00fe Kc 2 vcPM \u22c5ec2M \u22c5ec2M \u00f048\u00de q \u00bc \u2212\u03c9P\u22c5 d\u03c9P M d\u03c6P rcM\u2212\u03c9 P M vcM \u00fe \u03c92 P\u22c5 d2aPcM d\u03c62 P \u00f049\u00de The vectors in Eqs. (47)\u2013(49) are all expressed in the coordinate system oM \u2212 xMyMzM. The unit normal vector of the pinion cutter is ncM , v cP M is the relative velocity vector between the cutter and the pinion, \u03c9cP M is the relative angular velocity vector between the cutter and the pinion, Kc 1 is the principal curvature in the first main direction ec1 of the point on the pinion cutter, Kc 2 is the principal curvature in the second main direction ec2 of the point on the pinion cutter, e c 1M is the vector obtained by expressing ec1 in the coordinate system oM \u2212 xMyMzM, ec2M is the vector obtained by expressing ec2 in the coordinate system oM \u2212 xMyMzM, and \u03c9P M is the angular velocity vector of the pinion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003043_ssi52265.2021.9467023-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003043_ssi52265.2021.9467023-Figure12-1.png",
        "caption": "Fig. 12. A proximity sensor is integrated in 3D printed housing that is attached to an adaptive 3-finger gripper by Robotiq. The correlation between the distances measured by the proximity sensor and the position increments by the robot gripper is almost linear.",
        "texts": [
            " The algorithm for distance determination described above was applied. Fig. 11 shows echo signals from a planar reflector. A reliable detection with a constant gain was possible between 6 mm and 23 mm. Using TGC, the signal could still be detected at a distance of almost 40 mm (not shown). This demonstrates the robustness of the algorithm used for highly noisy signals. A single channel proximity sensor was finally assembled in a 3D-printed housing and attached to an adaptive 3-finger gripper by the company Robotiq (see Fig. 12). Measurements of the reflection from a planar surface were successfully recorded with the sensor driven with pulse sequence B. This setup allows the detection of objects up to 10 mm distance with a high spatial resolution and repeatability. A dynamic distance measurement was challenging, due to the narrow directivity pattern of the sensor. A multi-channel sensor array and/or a broader sound field would be more tolerant to tilt and rapid movement in the future. V. CONCLUSION A novel technology and sensor platform for the on-chip integration of tactile and proximity sensing modality based on CMUTs was developed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000729_humanoids43949.2019.9035004-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000729_humanoids43949.2019.9035004-Figure5-1.png",
        "caption": "Fig. 5. Exploded schematic of the MEMS RPA. Left: RPA housing layers; each layer contains a set of three deflection springs to align and secure an electrode. Upper right: electrode grid set; the edge of each grid has three notches that mate with the spring tips of the corresponding housing layer. Lower right: assembled RPA and cross-section.",
        "texts": [
            " The MEMS RPA design makes use of electrodes with apertures that share a common packing density with different aperture sizes to mitigate the effects of ion ray interception found in the experimental characterization of the hybrid RPA using an ion source. Inter-electrode aperture alignment was enforced via precise machining and layout features etched in the grids that mate with the tips of a set of MEMS deflection springs [27] monolithically attached to the housing that improves the alignment precision by an order of magnitude [28] compared to the hybrid RPA design (Fig. 5). Similar to the hybrid RPA design, the electrodes are gripped from a few points and no dielectric is present near the flow of charged species to prevent charging of non-conductive surfaces and to mitigate shorting via surface breakdown. The sensing area of the probe is a circle with 7.5 mm diameter and the inter-electrode spacing is 200 \u03bcm, which is the difference in thickness between the grid and corresponding housing layer. The electrodes are assembled to the housing using curved compliant structures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002216_s12541-020-00457-y-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002216_s12541-020-00457-y-Figure11-1.png",
        "caption": "Fig. 11 Six-bar linkage of proposed FCU",
        "texts": [
            " Clamping range of A2 is 73% larger than that of A1 and 44% larger than that of A3. This result is before the dimension synthesis of each alternative, thus the possibility remains that the difference will become even larger after dimension synthesis. This proves that the unique characteristics of A2 are also effective in compliant linkage analysis. 1 3 An FCU device was constructed to test its performance. The overall design of the final FCU is shown in Fig.\u00a010 and the internal linkage structure is presented in Fig.\u00a011. The device has an electric motor as an actuator with a maximum torque of 0.2\u00a0N\u00b7m. Three staged spur gears were applied to amplify the torque from the motor; the input torque of link 1 was 20\u00a0N\u00b7m. Link 1 acts as both a gear and a linkage. Link 2 is an L-shaped link with a slider which has space in the middle to enable link 1 and link 3 to pass through when the FCU unclamps. Link 3 is the compliant link which was motivated by a wave spring [28]. Link 4 is the final link that transfers input torque to output torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002060_0959651820977572-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002060_0959651820977572-Figure1-1.png",
        "caption": "Figure 1. The overall structure of VGMSFW.",
        "texts": [
            " Force analysis and modeling of the conical MB-rotor system have been reported in section \u2018\u2018System modeling.\u2019\u2019 Section \u2018\u2018Design of PIDNN Control\u2019\u2019 describes the method of PIDNN control method in detail. Section \u2018\u2018Simulation research\u2019\u2019 gives the comparison results between the proposed algorithm and other algorithms. Section \u2018\u2018Experimental investigations\u2019\u2019 verifies the effectiveness of the proposed algorithm. Finally, the article is concluded in section \u2018\u2018Conclusion.\u2019\u2019 The overall structure of VGMSFW is shown in Figure 1. The system consists of one base, one shaft, one rotor, one motor, one conical MB, two pairs of Lorentz force MBs, two touchdown bearings and displacement sensors. The motor drives the rotor to rotate at high speed. The conical MB controls the threetranslation DOF of the rotor, and the Lorentz MB controls the two-tilt DOF of the rotor. The displacement and tilt angle of the rotor are measured by displacement sensors. When the MB fails, the touchdown bearings are used to ensure that the high speed magnetically suspended rotor slows down safely"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001637_icarm49381.2020.9195279-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001637_icarm49381.2020.9195279-Figure5-1.png",
        "caption": "Fig. 5. Double inverted pendulum model. (a) Roll in coronal plane. (b) Pitch in sagittal plane.",
        "texts": [
            " The desired torques of double ankles in both pitch and roll directions are also adjusted by this controller. In this section, we first introduce the definition and dynamics of the double inverted pendulum, and then explicates the design of the torso posture controller which maintains the balance and calculate the desired torques, and the realization about the dynamic torso movement. In coronal (roll) and sagittal (pitch) planes, the humanoid robot in standing is simplified as two double inverted pendulums as shown in Fig. 5, which means we control the robot in roll and pitch orientations separately. In the following, we will introduce the model and controller in one plane, which is the same as the other plane. As shown in Fig. 5, the double inverted pendulum contains two joints and two poles. The first joint locates in the middle of left and right ankles, and the second joint is in the middle of left and right hips. The first pole, which is also the lower pole, is set from the first joint to the second joint. And the second pole is fixed on the torso, which has the same postures with the upper body of the robot. Tab. I gives the symbolized parameters of the double inverted pendulum, and we will use g for the gravity. Then the relationship between the torso posture qT and the double inverted pendulum joint angles q1, q2 can be presented as qT = q1 + q2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000422_lra.2019.2961302-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000422_lra.2019.2961302-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_63_0002060_0959651820977572-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002060_0959651820977572-Figure2-1.png",
        "caption": "Figure 2. Force analysis of the conical MB rotor system: (a) the forces on the tilted rotor and (b) partial enlarged detail of gap.",
        "texts": [
            " The system consists of one base, one shaft, one rotor, one motor, one conical MB, two pairs of Lorentz force MBs, two touchdown bearings and displacement sensors. The motor drives the rotor to rotate at high speed. The conical MB controls the threetranslation DOF of the rotor, and the Lorentz MB controls the two-tilt DOF of the rotor. The displacement and tilt angle of the rotor are measured by displacement sensors. When the MB fails, the touchdown bearings are used to ensure that the high speed magnetically suspended rotor slows down safely. According to Xiang and Tang,6 the forces are shown in Figure 2(a) when the rotor is tilted. Frx+ and Frx are suspension forces generated by the radial MB in the positive and negative directions of X-axis, respectively. Fux+ and Fux represent magnetic forces at the upper end of the conical MB, and Fdx+ and Fdx represent the ones at the lower end. The summation of these four forces equals to the gravitation of rotor. FL is the magnetic pull force generated by the Lorentz MB on the rotor tilt, b is the angle between the conical surface of the stator and Z-axis, and b=458",
            " The space rectangular coordinate system is established with the geometric center of the rotor as the origin, and the dynamic model of the 5-DOF rotor is obtained as m\u20acdx =Frx+ Frx + Fux +Fdx Fux+ Fdx+\u00f0 \u00de cosb m\u20acdy=Fry+ Fry + Fuy +Fdx Fuy+ Fdy+ cosb m\u20acdz = Fdx+ +Fdx Fux+ Fux \u00f0 \u00de sinb mg Jx \u20acu+ JzO _c=Tx Jy\u20acc JzO _u=Ty 8>>>>< >>>>: \u00f01\u00de When the rotor is tilted by an angle u around Yaxis, u represents the angle between Fux+ and Z-axis, and u+ u \u2019 458. The rotor\u2019s translational control is disturbed by the radial and axial component forces of the conical MB, and the magnetic forces at the conical gap are projected onto X-axis and Z-axis as follows Fcx= ffiffiffi 2 p 2 Fux Fux+ +Fdx Fdx+\u00f0 \u00de cosu + ffiffiffi 2 p 2 Fux +Fux+ Fdx Fdx+\u00f0 \u00de sinu \u00f02\u00de Fczx= ffiffiffi 2 p 2 Fdx +Fdx+ Fux Fux+\u00f0 \u00de cosu + ffiffiffi 2 p 2 Fdx Fdx+ +Fux Fux+\u00f0 \u00de sinu \u00f03\u00de Figure 2(b) shows the partial enlarged detail of gap around Fux+ . The radial and axial displacements at the center of the cone surface of the tilted rotor are Ddx and Ddz, respectively. When Ddx \u2019La sinu and Ddz \u2019Lr sinu, the variation of the air gap is Dd= ffiffiffi 2 p Ddz Ddx\u00f0 \u00de 2 = ffiffiffi 2 p Lr La\u00f0 \u00de sinu 2 \u00f04\u00de Lr and La are the distances from the center of rotor\u2019s cone-shaped surface to Z-axis and X-axis, respectively. According to the design size of the conical MB, Lr=47:82mm and La =34:57mm. When the rotor is tilted by 18, Dd is about 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002253_cac51589.2020.9327623-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002253_cac51589.2020.9327623-Figure1-1.png",
        "caption": "Figure 1: Quadrotor UAV body diagram.",
        "texts": [],
        "surrounding_texts": [
            "Considering a multi-UAV system with one leader labeled 0 and N followers labeled 1, 2, . . . , N . The interaction topology among the N followers can be described by a weighted directed graph G = (W , E ,A), where W = {w1, w2, . . . , wN} denotes the set of nodes, E = {eij = (wj , wi), wi, wj \u2208 W} represents the set of edges, and A = [aij ]N\u00d7N is the weighted adjacency matrix with i, j \u2208 {1, 2, . . . , N}. In addition, eij denotes the edge formed by nodes wj and wi, where wj and wi are called the parent node and child node, respectively. Moreover, aij > 0 represents the weight of edge eij if eij \u2208 E , and aij = 0 if not. Besides, one assumes that aii = 0, \u2200i = 1, 2, . . . , N . A directed path between nodes wi and wj is defined by a series of edges (wi, wi1), (wi1, wi2), . . . , (wil, wj), where wik (k = 1, 2, . . . , l) are different nodes of the graph. For a directed graph G, if there exists at least one node that has directed paths to all the other nodes, then it is said to have a directed spanning tree. The in-degree of node wi is defined as degin(wi) = \u2211N j=1,j =i aij . Then the indegree matrix D and the Laplacian matrix L are defined as D = diag(degin(wi), i = 1, 2, . . . , N) and L = D \u2212 A, respectively. Assume that the communication between the leader and followers is unidirectional, which means that the followers can get the status of the leader, but otherwise it is not. The interaction weight between the leader and follower i is denoted by ai0. ai0 > 0 if the follower i can get the status of the leader, and ai0 = 0 if not. In addition, denote H = diag(a10, a20, . . . , aN0) and LH = L+H . Lemma 1. If the directed graph G contains a directed spanning tree from the leader, then the matrix LH is invertible."
        ]
    },
    {
        "image_filename": "designv11_63_0000116_012069-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000116_012069-Figure4-1.png",
        "caption": "Figure 4.Free Body Diagram of VTOL System",
        "texts": [
            " Equation (1) shows the voltage-current relationship of the DC motor actuator, dt di LiRv m mmmm (1) The voltage across the motor is vm, im current drawn by the motor with resistance Rmand inductance of the motor is Lm. Equation (2) shows the voltage-current relation in the transfer function form, mm m m RsL V I (2) The transfer function of the system has Vm i.e. voltage induced to the motor as the input to the actuator and Im i.e. the output current to be induced to the system as output. Unlike the DC motor, the VTOL system can be characterized by a second order equation. The VTOL system consists of propeller actuator, VTOL body and counter weight. Fig 4 shows the free body ICMSMT 2019 IOP Conf. Series: Materials Science and Engineering 561 (2019) 012069 IOP Publishing doi:10.1088/1757-899X/561/1/012069 diagram depicting these parts. Equation (3) gives the transfer function of the VTOL system [1] which is derived from the free body diagram. J K s J B sJ K sP t 2 )( (3) where, Kt-thrust current torque constant Nm/A J -moment of inertia in kgm2 B -viscous damping in Nm/(rad/s) K-stiffness of the VTOL system in Nm/rad Referring [1] and [2] the parameters of the transfer function of the VTOL system were found as in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002755_tmag.2021.3074935-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002755_tmag.2021.3074935-Figure3-1.png",
        "caption": "Fig. 3. No load flux distributions of Models 1\u20134 when dc current is 13.44 A. (a) Model 1: 8/6/2. (b) Model 2: 5/3/2. (c) Model 3: 7/3/2. (d) Model 4:7/6/1.",
        "texts": [
            " 2(b)\u2013(d), the phase currents are i A+ = \u221a 2Iac sin(\u03c9et + \u03b1)+ Idc i A\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1)\u2212 Idc iB+ = \u221a 2Iac sin(\u03c9et + \u03b1 \u2212 2\u03c0/3)+ Idc iB\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1 \u2212 2\u03c0/3)\u2212 Idc iC+ = \u221a 2Iac sin(\u03c9et + \u03b1 + 2\u03c0/3)+ Idc iC\u2212 = \u221a 2Iac sin(\u03c9et + \u03b1 + 2\u03c0/3)\u2212 Idc (3) where Idc is the dc component, Iac is the ac component, and \u03b1 is the phase A current angle. Moreover, two main drive circuits in Fig. 2(a) and (b) can output the required currents of (2) and (3), respectively. When only dc current is injected into windings with described current direction in (2) and (3), an exciting field will be produced in the air gap with the rotor permeance modulation effect. No load flux distributions of Model 1\u20134 with a dc current of 13.44 A are observed when the flux linkage of Phase A is maximum, as shown in Fig. 3. And the radial air gap flux density waveforms are exported and shown in Fig. 4. It can be seen that the magnetic flux distribution of Model 1 is centrosymmetric while that of Models 2 and 3 are asymmetric, because Nr and Pdc of Models 2 and Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 24,2021 at 17:02:44 UTC from IEEE Xplore. Restrictions apply. 3 are both odd while Nr and Pdc of Model 1 are both even. For Model 4, Nr is odd and Pdc is even, so its flux distribution is symmetric"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001378_s11668-020-00961-3-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001378_s11668-020-00961-3-Figure2-1.png",
        "caption": "Fig. 2 Eye-end composite leaf spring",
        "texts": [
            "com service condition of such spring is relatively good, and there is no need to specially design the joint. The longitudinal composite leaf spring is generally installed on the vehicle with a non-independent suspension. The service condition is relatively harsh, and a joint with reliable performance needs to be designed. According to the type of connection with the vehicle body, longitudinal composite leaf springs can be divided into eye-end leaf springs and sliding leaf springs. The common structure of the eye-end leaf spring is shown in Fig. 2. Both ends of the leaf spring are metal eye joints, the leaf spring body and the eye joints are connected by bolts. Two metal plates are attached to the middle of the leaf spring body for eliminating the stress concentration when the U-bolt is clamped. The structure of the sliding leaf spring is shown in Fig. 3. The middle structure of the sliding leaf spring is basically as same as that of the eye-end leaf spring. On both ends of the leaf spring are gaskets and hooks, usually made of metal. The gaskets are mainly responsible for the force transmission between the spring body and the support, and the hooks can prevent the leaf spring from leaving the suspension"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002419_iros45743.2020.9341316-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002419_iros45743.2020.9341316-Figure2-1.png",
        "caption": "Fig. 2. A rigid object grasped by environment and manipulators.",
        "texts": [
            " The reaction forces produced at the contact along the edge can be used to partly balance the weight of the object. This, in turn, reduces the force required to grasp the object against the effect of its own weight. We have evaluated our method in simulation, for (1) tilting a large hollow cylindrical object using a parallel jaw gripper, and (2) tilting a box about its edge using two arms of the Baxter robot. Consider a rigid object which is in contact with the environment at m points and grasped by n manipulators at n positions as shown in Fig. 2. Contact coordinate frames {Ci} and {Ej} (where i = 1, .., n and j = 1, ...,m) are attached to the object at each manipulator and environment contact, respectively, such that n-axis of the frames is normal (inward) to the object surface and two other axes, t and o, are tangent to the surface. Positions of the contact frames in the inertial coordinate system (X ,Y ,Z) are represented by pCi \u2208 R3 and pEj \u2208 R3. A contact model can be used to impose constraints on the forces that arise at the contact locations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001115_icieam48468.2020.9111886-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001115_icieam48468.2020.9111886-Figure3-1.png",
        "caption": "Fig. 3. Cross-section of the elastic joint",
        "texts": [
            " This work is devoted to the development of control system of a single elastic joint, which was previously presented in [19]. The structure of the controller is based on the ideas of [20], specifically; the motor variable substitution is applied to build an equivalent mathematical model of the actuator, introducing a virtual damping part. An algorithm to tune the control parameters is hereby contributed as elaboration of the basic approach. III. JOINT DYNAMIC MODEL The joint [19], which is shown in Fig. 3, comprises three main components: a brushless DC-motor RobodriveILM70x18, a wave gear Harmonic Drive CobaltLine-32-CPM, and a custom elastic element with stiffness from 500 to The dynamic model of the joint is implemented in MATLAB-Simulink and contains three basic blocks corresponding to the three basic parts of the joint, see Fig. 4. Authorized licensed use limited to: Imperial College London. Downloaded on June 13,2020 at 14:23:50 UTC from IEEE Xplore. Restrictions apply. The motor is modeled traditionally as DC-motor with motor voltage and load torque as inputs and motor speed and position as outputs, see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002250_s12666-020-02152-y-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002250_s12666-020-02152-y-Figure6-1.png",
        "caption": "Fig. 6 Coupled simulation of mold filling and casting solidification",
        "texts": [
            " Explicit time integration scheme for transient simulation was employed, and the time-step for solver computations was automatically selected based on stability criteria. The metal pouring temperature is one of the inputs for mold filling simulation. The computed temperature at the end of filling serves as the initial temperature for solidification simulation. In 2011, the NIIST team joined hands with IIT Bombay and an industry partner (3D Foundry Tech) to develop an advanced simulation module called FLOW? with improved features, user interface and computational performance (Fig. 6). The default inputs including thermo-physical properties and interface heat transfer coefficients are taken from a comprehensive database to minimize user inputs [10]. The mesh generation is also automatic, though the user can control the desired speed and accuracy of results. The software computes mold filling velocity, liquid fraction, solidification temperature and cooling rate at various points inside the casting. These computations take only a few hours for even large and complex castings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001377_14763141.2020.1787498-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001377_14763141.2020.1787498-Figure1-1.png",
        "caption": "Figure 1. Definition of joint angles.",
        "texts": [
            " We defined the x-axis as the frontal axis of the lifter and the y-axis as the longitudinal axis of the lifter. The position of the barbell at the lift-off was set as the coordinates\u2019 origin, and the direction of the lifter\u2019s line of sight was set at the positive quadrant of the x-axis. The centre of mass (COM) of the lifter during the snatch motion was calculated based on body segment inertial parameters (Yokozawa et al., 2016) to examine the lifter\u2019s motion. To examine the joint movement during snatch, we evaluated the hip joint, knee joint, ankle joint, trunk, and trunk-arm angles (Figure 1). Joint angles were calculated as the angle between the two segments on either side of the joint that projected in the sagittal plane. The trunk angle was the absolute angle of the trunk segment (vector of the hip joint process to the shoulder joint) on the x-axis. Increases and decreases in the trunk angle indicated backward and forward bending, respectively. The trunk angle is 90\u00b0 when the trunk segment and x-axis are perpendicular. Trunk-arm angle was formed by a vector connecting the shoulder joint process to the barbell and another connecting the shoulder joint process to the hip joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000922_pedstc49159.2020.9088409-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000922_pedstc49159.2020.9088409-Figure6-1.png",
        "caption": "Fig. 6. 2-D cross-section view of a 6/7 VFRM",
        "texts": [
            " Windings and their connection are significant factor in VFRMs as individual phase coils should be connected in opposite directions when the number of rotor poles is odd. Fig. 5 shows the vectors of back-EMF in these ratios of rotor and stator teeth number. (a) (b) (a) (b) (c) (d) III. WINDING FUNCTION MODELING PROCEDURE OF A 6/7 VFRM In this section, the basic principal of WFM is going to be introduced. Turn, winding, and airgap function play key roles in winding function modeling of VFRMs. Turn function proposes the placement of windings among machine geometry. A 2-D cross-section view of a 6/7 VFRM is shown in Fig. 6. The windings of Phase a, phase b, and phase c, are in red, green, and blue, respectively. Also, the DC field windings are in yellow. . Authorized licensed use limited to: Auckland University of Technology. Downloaded on May 29,2020 at 07:39:29 UTC from IEEE Xplore. Restrictions apply. The relation between turn function and winding function is as below: ( ) = ( ) \u2212 < ( ) > (1) where ( ) , ( ) , and < ( ) > are winding function, turn function, and mean of turn function, respectively [18]. The winding functions of abovementioned machine are shown in Fig 7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001218_ec-10-2019-0447-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001218_ec-10-2019-0447-Figure4-1.png",
        "caption": "Figure 4. Computed temperature field during the deposition of the fourth layer for Vs = 200mm/min and Plaser = 600W",
        "texts": [
            " Indeed, the flow in the melt improves the mixing thereof and reduces the temperature compared to the purely thermal calculation of Peyre et al. (2017), which overestimates the temperature as reported by Manvatkar et al. (2015) and Khairallah et al. (2016). L2 error norms are calculated for comparing the experimental temperatures with the calculated ones. For Plaser = 400W, simulation results obtained by (Peyre et al., 2017) (L2 = 644.2) are better than our results (L2 = 937.9); but for Plaser = 600W, our results (L2 = 938.1) are better than those of (Peyre et al., 2017) (L2 = 1,192.5). Figure 4 shows the temperature distribution obtained by our model during the deposition of the fourth layer for Plaser = 600W. The shape of the isotherms is not symmetrical with Table 1. Main material properties used in the simulations Solidus temperature, Ts (K) 1,873 Liquidus temperature, TL (K) 1,923 Density at TL, rL (kg/m3) 3,800 Latent heat of fusion, DHm (kJ/kg) 2906 5 Specific heat capacity, Cp (J/kg/K) 1126 Thermal conductivity l (T) (W/m/K) 0.32\u00fe 1.46 10\u20132T 1400< T< 1,850K 6.66\u00fe 1.83 10\u20132T 1950< T< 2,700K Emissivity, \u00ab (T) 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure13-1.png",
        "caption": "Fig. 13. Demagnetization ratio of STPSPM machines with eccentricities at 16ms during 3PSC. (a) Static eccentricity, 10-pole/12-slot. (b) Rotating eccentricity, 10-pole/12-slot. (c) Static eccentricity, 10-pole/9-slot. (d) Rotating eccentricity, 10-pole/9-slot.",
        "texts": [
            " (a) (b) -40 -20 0 20 40 0 4 8 12 16 20 24 C u rr e n t (A ) Time (ms) C u rr e n t (A ) A B C Static eccentricity No eccentricity -40 -20 0 20 40 0 4 8 12 16 20 24 C u rr e n t (A ) Time (ms) Rotating eccentricity No eccentricity B C A Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. The demagnetization distributions due to 3PSC are shown in Fig. 13. For symmetrical SPM machines, when there is no eccentricity, the demagnetization distribution is rotationally symmetrical, as shown in Fig. 5(a). However, with static or rotating eccentricity, the demagnetization is not rotationally symmetrical and gathers around the PMs facing larger air-gap length. This can be explained by open-circuit and armature field working points, and the influence of static eccentricity on symmetrical STPSPM machine is analyzed as an example. It is defined that positive x direction in the stator reference is 0 mechanical degrees whilst the counter-clockwise direction is the positive direction, as shown in Fig",
            " On the contrary, at the position of 330 mechanical degrees, the open circuit flux density increases from -0.79T to -0.87T whilst the armature field flux density increases from 0.27T to 0.3T. The resultant flux density increases from - 0.52T to -0.57T and thus demagnetization is much mitigated on the mid-PM at such position due to smaller airgap length when static eccentricity is applied. Therefore, the demagnetization generally tends to gather to PMs facing larger airgap length. In addition, the conclusions are also applicable to asymmetric STPSPM machines, as shown in Fig. 13. (a) -1 -0.5 0 0.5 1 1.5 0 60 120 180 240 300 360 F lu x d e n s it y ( T ) Position (Mech. Deg.) no eccentricity static eccentricity Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. For the 10-pole/12-slot STPSPM machine, iq=8A is again applied for a mechanical period after 3PSC ends to obtain the post-demagnetization UMF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003203_s42835-021-00864-9-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003203_s42835-021-00864-9-Figure7-1.png",
        "caption": "Fig. 7 Illustration of receiving coil misalignment direction (a) x-direction misalignment, (b) y-direction misalignment",
        "texts": [
            " The induced voltage is defined as follows: Similarly, the inductance voltage u2, u3, and u4 of coil C2, C3 and C4 can be obtained. Figure\u00a06 shows the four-coil voltage comparison circuit, and Table\u00a01 shows the different results that may occur. When the voltage difference \u2206u12 is less than zero or \u2206u34 is greater than zero, it means that the coil is shifted to the positive direction of the coordinate axis. On the contrary, it means that the coil is shifted to the negative direction of the coordinate axis, Fig.\u00a07 show the schematic diagram of the offset direction of the receiving coil. (16)u1 = u01 25000 \u22c5 R(1 + R15 R16 ) 1 3 In this section, the effects of the structural parameters, such as shapes and dimensions of the receiving coil is investigated, and the electromagnetic shielding structure is proposed to strengthen the coupling and lower the cost. The analysis of the receiving coil under parameters variations are provided. In this paper, the receiving coil adopts four coils in series connection. According to the connection modes of the four coils, it can be divided into the surface intersecting structure and the edge intersecting structure, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure45.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure45.1-1.png",
        "caption": "Fig. 45.1 Semi-circular bucket-type metering mechanism attached with a chain",
        "texts": [
            " The research focused on the design, development, and testing of self-propelled mini ginger planter suitable for the hilly region of northeast India, where the use of the large size of machinery is not feasible. Ginger is irregular in shape and planted by cutting the portion randomly with buds of varying lengths from 20 to 30 mm. The length of ginger used for plantation is 2.5 to 30 mm in length, the thickness of 19.23 mm with a standard deviation of \u00b11.94 mm [1], and width was taken to be the maximum length of the ginger seed. For picking the irregular shape, the chain and bucket metering mechanism was used for the planter, as shown in Fig. 45.1. The overall length and width of the cup were kept at 45 mm. The depth was decided as 22 mm to avoid the chances of coming two seeds at a time during operation. As the bucket passes through the hopper filled with ginger seed, the bucket picks up the seed from the hopper. The bucket inverts as they pass around the top sprocket, and the seeds pass through the discharged tube. The whole pathway is enclosed with a rubber material to avoid damage of seed until it gets delivered in the furrow opener through the discharged tube"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure8-1.png",
        "caption": "FIGURE 8. Displacement of modified frame",
        "texts": [
            " According to result in the Figure 5 and Figure 6, the frame must be modified to enhance its performance since the structure can be considered as fail. The value of intrusion and also the stress were not fulfilling the requirement of ECE 66 standard. Some part of the frame has been modified by increasing its thickness to reduce the stress and also the value of intrusion. List of modification was tabulated in Table 3. The location of modified part in the frame were depicted on Figure 7. 030153-5 The simulation was done for the modified frame and the result for displacement, stress and also intrusion were depicted in Figure 8, Figure 9 and Figure 10 respectively. Based on the result of simulation, the displacement of frame was decrease. The magnitude of maximum deflection is 359.4 mm, and minimum displacement is 39.94 mm. 030153-6 The von Misses stress actually increase compared to the existing frame, but during the rolling accident, the collapse part is needed to absorb the impact energy from the external load. Therefore, the passenger is safe since they received the minimum impact energy. The maximum von Misses stress is 230"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000922_pedstc49159.2020.9088409-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000922_pedstc49159.2020.9088409-Figure14-1.png",
        "caption": "Fig. 14. Distribution of the magnetic flux density of the proposed 6/7 VFRM",
        "texts": [
            " As it can be seen, the armature mutual-inductances is constant while the armature and field winding mutual inductances has a sinusoidal profile. Therefore, the armature and field mutual inductances variation, leads to the electromagnetic torque production. Fig. 13 shows the flux lines of the proposed Authorized licensed use limited to: Auckland University of Technology. Downloaded on May 29,2020 at 07:39:29 UTC from IEEE Xplore. Restrictions apply. VFRM, which is the result of FEM. Also, distribution of the magnetic flux density of the proposed VFRM is investigated in Fig. 14. It can be seen that alignment of rotor pole and stator tooth, leads to the maximum flux density which is 1.5T. The torque capability of VFRM is computed by WFM and the results are compared with the results of experimental test. Fig. 15 (a) and Fig. 15 (b) show the average torque which is computed by WFM and meaured by experimental tests in several q-axis currents with field current of 1 A and 2 A, respectively. It can be seen that the error of resultant average torque from WFM and experimental is less than 10 %, which proves the accuracy of the proposed WFM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001018_isef45929.2019.9097061-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001018_isef45929.2019.9097061-Figure4-1.png",
        "caption": "Fig. 4. Evaluation result of minimum torque.",
        "texts": [
            " This is because the magnetic flux for two coils passes through the pole in the conventional model, whereas in the proposed model the magnetic flux for only one coil passes through the pole. IV. OPTIMIZATION OF POLE ARRANGEMENT AND TORQUE In this section, the proposed stator pole arrangement with 5 phase and 10 coils is optimized. A genetic algorithm is used for a pole arrangement, and the evaluation is performed by the torque area method [4]. By using the torque area method, torques around every axes can be evaluated. The evaluation results of the optimized model using the conventional and proposed pole structures (models) are shown in Fig. 4. These maps of minimum torques are obtained by changing the rotor inclination angle from 0 deg. to 40 deg. in the latitudinal direction and from 0 deg. to 180 deg. in the longitudinal direction. As a result of the conventional pole structure, the average torque is approximately 0.28 Nm. The average torque of the proposed model is about 0.52 Nm, which is approximately 87% higher than the conventional model. On the other hand, torque area where the minimum torque is less than 0.01 Nm can be seen in the proposed model as well as the conventional model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002794_j.ijfatigue.2021.106312-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002794_j.ijfatigue.2021.106312-Figure2-1.png",
        "caption": "Fig. 2. Box-shaped filler-reinforced natural rubber gearbox mount [22].",
        "texts": [
            " This mount is Nomenclature Anqn ,Bnqn ,Cxqn coefficients of the gain functions bn damping parameters of the several model paths cn stiffness parameters of the several model paths czprog negative progressive stiffness for a displacement value of qz = 11 mm in the z-direction as a function of the transverse displacements qx and qy fn(qn) gain functions of the several model paths F1 mount force of the generalised mount main direction F1n forces of the individual model paths of the main mount direction FHo force calculation upper hysteresis section FHu force calculation lower hysteresis section FHyst hysteresis force adapted model according to Dronka & Rauh Freal measured operating loads Fsim simulated operating loads FUP force value last hysteresis reversal point Fx,y,z mount forces (x-,y-,z-direction) k Slope of Wo\u0308hler curve kf ,kx parameter hysteresis element adapted model according to Dronka&Rauh LWz loss work of the z-direction qn generalised excitation q\u0307n generalised excitation velocities qUP displacement value last hysteresis reversal point qx,y,z translational mount displacements (x-, y-, z-direction) Smeas signal damage measured operating loads Srel relative signal damage Ssim signal damage simulated operating loads x,y, z indices for the description of translational spatial directions S. Ernst et al. International Journal of Fatigue 150 (2021) 106312 installed in a small vehicle with a front transverse engine and front wheel drive (see Fig. 2). The mount has a complex structure, which is divided into a housing, basic load spring and several stops to support high forces in x, y and z direction. The measurement of the gearbox mount is performed on a 4-DOF elastomer test rig (see Fig. 3). With the test system it is possible to simultaneously apply three axial displacments and one rotational excitation to the components under optional additional thermal load. The test rig is designed to display stochastic signals with a maximum frequency content of approx",
            " In addition to these studies, a standard characterisation (measurement of quasi-static stiffness, dynamic stiffness and loss angle under uniaxial loading) is carried out for the mount. The displacement measurement on the test system is not performed directly on the adapter of the component, instead it is calculated via the movement of the hydraulic cylinders. At high forces, the stiffness of the test rig as well as the stiffness of the test adapters must be considered for an exact determination of the component displacement (displacement between the housing and the gearbox connetion, see Fig. 2). For this purpose the system has a compensation mechanism for each of the main axes. The stiffness compensation is applied for all displacement values determined on the test system. As already mentioned in the introduction, the operating loads required for the design and validation process of powertrain mounts are determined on the basis of real track tests. The track data used here consists of a section to evaluate driving dynamics and a section representing rough road conditions. The track is selected by the car manufacturer in such a way that a large spectrum of customer-relevant loads is represented",
            " In the last 28% of the signal, the higher frequency loads are shown, which are caused by the rough road section. The force\u2013time signals of the mount serve as the basis for the subsequent fundamental tests to investigate the change in uniaxial quasi-static mount properties (e.g. global stiffness and hysteresis) due to spatial loads. Within the context of these analyses, only measured values of the vertical mount direction (z-direction) are shown. The definition of the z-direction of the mount can be taken from Fig. 2. In contrast to the common practice for metallic materials, which are based on the component stress at local critical damage locations, the damage analysis for the gearbox mount is carried out based on the global force responses of the in the individual spatial directions. The method presented below is a state of the art procedure in the automotive industry for evaluating the load-time series, which serve as the initial values for the durability verification in the operating load test. Fig. 5a shows the normalized operating forces measured on the test rig for the zdirection in detail"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001375_speedam48782.2020.9161929-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001375_speedam48782.2020.9161929-Figure1-1.png",
        "caption": "Fig. 1 \u2013 The filling of flux barrier in PMaSRM: the traditional approach with regular sintered magnets",
        "texts": [
            " AIM OF THE WORK In recent times, different types of bonded magnets have been investigated and adopted in various electrical machine applications. Some prototypes of assisted reluctance machine both with NdFeB and ferrite bonded magnets have been prepared with the adoption of different manufacturing techniques [31]. In the proposed work the two different concepts concerning the use of ferrites magnets in assisted reluctance machine have been investigated. The first approach consists in the adoption of regular shape magnets, which take up a partial part of the volume of the flux barriers, as shown in Fig. 1. In the other case, the flux barriers are completely filled with bonded magnets, as reported in Fig. 2. The study began with the investigation of a particular SRM with a rotor geometry designed for the propose of lowering torque ripple [36-38]. The flux barriers shapes have been optimized by means of a multi-objected differential evolution (DE) algorithm [39-42]. Some geometric dimensions and rated parameters concern SRM are shown in TABLE I. 978-1-7281-7019-0/20/$31.00 \u00a92020 IEEE 670 Authorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on August 12,2020 at 02:43:59 UTC from IEEE Xplore. Restrictions apply. The obtained rotor geometry illustrated in Fig. 1 and Fig. 2, has been adopted for evaluations performed by means of finite element analysis (FEA). During FEA estimations three different cases have been compared: SRM (without permanent magnets), PMaSRM with parallelepiped shape ferrite sintered magnets and PMaSRM with ferrite bonded magnets. After careful simulation analysis, the appropriate prototype will be prepared. All proposed magnets have been characterized in the laboratory. Moreover, the involved prototype has been also built and tested in the laboratory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure4-1.png",
        "caption": "Fig. 4. Angles between the pendulum axis and coordinate axes of the CCS.",
        "texts": [
            " According to the principle of inertia ellipsoid [23], the relation between the inertia tensor (Jx, Jy, Jz, Jxy, Jyz and Jzx) of measured body with respect to CCS and the MOI of measured body about a line passing through its COG can be expressed as: \u23a1 \u23a2 \u23a2 \u23a3 JP1 JP2 \u22ef JPN \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 l2 1 m2 1 n2 1 \u2212 2l1m1 \u2212 2m1n1 \u2212 2n1l1 l2 2 m2 2 n2 2 \u2212 2l2m2 \u2212 2m2n2 \u2212 2n2l2 \u22ef \u22ef \u22ef \u22ef \u22ef \u22ef l2 N m2 N n2 N \u2212 2lN mN \u2212 2mNnN \u2212 2nNlN \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 Jx Jy Jz Jxy Jyz Jzx \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (5a) Od Yd Zd Xd Zr Zc Yc Xc Zb Yb XbP2 Xr YrOc Ob(P1) Zg Og Xg Yg Or P3 Fig. 2. Relationship between five coordinate systems. T. Li et al. Measurement 174 (2021) 108956 li = cos\u03b1i, mi = cos\u03b2i, ni = cos\u03b3i (5b) where \u03b1i, \u03b2i, \u03b3i are angles between the pendulum axis and coordinate axes, OcXc, OcYc, OcZc, of CCS, respectively, as shown in Fig. 4. N is the total number of measurements. It is known from Eq. (5a) that to calculate inertia tensor of the measured body defined in CCS, only the MOI of the measured body along the pendulum axis and the direction cosine of the pendulum axis in CCS under 6 different suspended postures of the measured body, are required to be measured. However, in order to reduce measurement errors, 10\u201312 different suspended postures of the measured body are carried out usually and then using least squares methods to calculate the inertia tensor of the measured body defined in CCS [8]: J = ( ATA )\u2212 1ATJP (6) where, J = [ Jx Jy Jz Jxy Jyz Jzx ]T (7) JP = [ JP1 JP2 \u22ef JPN ] T (8) A = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 l2 1 m2 1 n2 1 \u2212 2l1m1 \u2212 2m1n1 \u2212 2n1l1 l2 2 m2 2 n2 2 \u2212 2l2m2 \u2212 2m2n2 \u2212 2n2l2 \u22ef \u22ef \u22ef \u22ef \u22ef \u22ef l2 N m2 N n2 N \u2212 2lNmN \u2212 2mNnN \u2212 2nNlN \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (9) It is worth noting that the condition number of A affects the calculated results of J",
            "00276s, eR=0.00031m . Taking a cylindrical body as studying example, the MOI around its central axis is 0.53 kg\u2219m2 estimated from CAD mode. The calculated maximum error dJ is equal to 0.005 kg\u2219m2 using Eq. (51) so that the maximum relative error of MOI of the cylindrical body, dJ/J, is 0.94%. Assuming the measured pendulum axis has an angle \u03b8 with the theoretical pendulum axis as shown in Fig. 7, the angle \u03b8 of pendulum axis will lead to errors in the estimated angles of \u03b1, \u03b2 and \u03b3 as shown in Fig. 4. Angular error \u03b8 results in the measurement errors of MOI. Assuming the maximum angular errors between measured pendulum axis and coordinate axes of the CCS are \u03b8 too. Hence the direction cosine of the measured pendulum axis in CCS is \u23a7 \u23a8 \u23a9 cos(\u03b1 + \u03b8) = cos\u03b1cos\u03b8 \u2212 sin\u03b1sin\u03b8 cos(\u03b2 + \u03b8) = cos\u03b2cos\u03b8 \u2212 sin\u03b2sin\u03b8 cos(\u03b3 + \u03b8) = cos\u03b3cos\u03b8 \u2212 sin\u03b3sin\u03b8 (52) If error of the measured axis, \u03b8, is small enough, Eq. (53) existed: cos\u03b8 \u2248 1, sin\u03b8 \u2248 \u03b8 (53) Therefore, the direction cosine of the measured axis is simplified as: \u23a7 \u23a8 \u23a9 cos(\u03b1 + \u03b8) \u2248 cos\u03b1 \u2212 \u03b8sin\u03b1 cos(\u03b2 + \u03b8) \u2248 cos\u03b2 \u2212 \u03b8sin\u03b2 cos(\u03b3 + \u03b8) \u2248 cos\u03b3 \u2212 \u03b8sin\u03b3 (54) The square of sin and cos functions in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002279_1.4006737-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002279_1.4006737-Figure16-1.png",
        "caption": "Fig. 16 Stretcher carrying on a stairway",
        "texts": [
            " Finally, the local coordinate frame of the patient RHumanPoint, as well as the stretcher coordinate frame, can be defined. Note that the above relation is effective even when the forward transporter is engaged in stairway walking while the other transporter is engaged in straight-line walking in a corridor. On the other hand, the distance between the transporters is given as a constant value L1 (Fig. 15). Drip injection behavior during stretcher transportation is considered by adding an injection-holding agent to the transporters, as shown in Fig. 16. The injection holder is located at an intermediate position between the two stretcher transporters. The holding agent also maintains an adequate distance from the stretcher. This agent behaves in the same manner as the other evacuee agents, except for the injection holding. Here, the injection tube length can also be considered as one of the simulation properties for patient transportation. In mixture evacuation simulations, all of the agents, such as patients along with their transporters, as well as nonhandicapped agents, move autonomously under predetermined behavior rules",
            " 2 were used, in the following, the behavior of KDH cooperative models in stairway walking is primarily discussed. The stairway model is 1.8 m in width, 0.2 m in height, and has eight steps. The patient transportation speed in the stairway is defined to be 50% of the evacuee agent speed used in Sec. 4. When L1\u00bc 2.2 m and the injection tube length is 2 m, three transporters carry the stretcher down a stairway while maintaining adequate separation for careful cooperation. The lead transporter, followed by the injection holder and the rear transporter move down stairway, as shown in Fig. 16. After the lead transporter reaches a landing, the stretcher, and then the rear transporter, reaches the landing. Meanwhile, the injection holder on the stairway comes to a halt in order to adjust the walking pace to be compatible with the other transporters. The lead transporter makes a wide turn in the landing and the rear transporter follows the leader with a slight time lag. Note that the injection holder can realize a cooperative motion naturally by making a small turn as well as adjusting his/her walking speed on the landing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001118_10426914.2020.1772489-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001118_10426914.2020.1772489-Figure4-1.png",
        "caption": "Figure 4. Construction of gear tooth profiles.",
        "texts": [
            " The same procedure is used to construct the profile of the coast side with various pressure angles. The base circles for the profiles and root circle are also drawn. A line is constructed from the involute profile limiting radius to a tangent point to the base circle. Two circles are drawn by considering line length as the radius for load side profile and host side profile of the adjacent tooth. A circle is constructed such a way that it is tangent to root circle and the two circles constructed as discussed above. The three circles and the tooth involutes profiles are shown in Fig. 4. The unwanted portions of the constructed lines are trimmed to get the complete gear tooth profile. The final profile is shown in Fig. 5. The gear parameters used in this study are given in Table 6. The AutoCAD DXF file is imported into the hyper mesh software. There the file is converted into neutral form (IGES). The IGES format of the profile has been imported to the machine to develop the WCEDM program. The gear blanks are case hardened and carburized with surface hardness value of 55HRC. Since the study of generated profile metrology is aimed at this work, only a few gear teeth were cut and inspected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000083_mees.2019.8896375-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000083_mees.2019.8896375-Figure2-1.png",
        "caption": "Fig. 2. Generator with magnetic gear.",
        "texts": [
            " CONSTRUCT DESCRIPTION The magnetic system of a multi-pole generator with an external rotor for a wind turbine is shown in Fig. 1. The external rotor of the generator consists of a outer core 1, on which the permanent magnets 2 of alternating polarity are fixed. The stator 3 of the electric generator has rectangular slots, on each of the stator teeth there is a separate winding coil, , the number of coils of each phase is 15; the number of turns of each coil - 10. The magnetic system of the electric generator with an integrated magnetic gearbox is shown in Fig. 2. The magnetic gearbox consists of an inner high-speed rotor 1, consisting of two layers of radially magnetized magnets 2, fixed segments of electrical laminated steel 4, an external low-speed rotor 5 with radially magnetized permanent magnets 6 mounted on a ferromagnetic ring core of structural steel, which is mechanically connected with a wind rotor shaft [1]. The stator 3 of the generator is completely identical in size to the stator of the generator with an external rotor, the number of phases is also equal to 3",
            " In this range, the maximum values of the electromagnetic torques for the external and internal rotors were determined. Then the next value of the height of the steel segments of the modulator was set, the dependence of the electromagnetic torques were calculated and the following maximum values of the torques were determined. The magnetic field picture of the generator with an external rotor is shown in the lower part of Fig.1. It should be noted that the elements of the generator magnetic circuit are not saturated. In the lower part of Fig. 2, the magnetic field of the GMG is shown, from which it is clear that the presence of a modulator leads to a complex redistribution of the magnetic field between the outer and inner rotors of the magnetic gearbox. Analysis of the results of calculations of this generator also confirmed the absence of saturation in the ferromagnetic elements of the magnetic circuit. With an increase in the height of the modulator segments, the magnitudes of the maximum torques of the outer and inner rotors increase, however, with a value of this height greater than 15 mm, the further increasing in the maximum value of torque is insignificant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001845_s00202-020-01132-1-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001845_s00202-020-01132-1-Figure8-1.png",
        "caption": "Fig. 8 The portrayal of the considered setup: a real housing,b simplified housing",
        "texts": [
            " Several different technical papers such as [3, 24] and [25] provide the typical form of the reduction in the velocity versus the distance from the fan forTEFCmachineswith varying ranges of power, which is operating at 50 Hz. Accordingly, Fig. 7 provides the velocity reduction versus the distance from the fan for 4-pole 10 kWTEFCmachinewith red dashed lines along with other machine graphs which were presented in [3, 24] and [25]. Figure 7 provides the starting value for the airflow in along the fin channels during the thermal analysis of TEFC machine. Figure 8a illustrates the real frameof themachine,which consists of semi-open fin channels and a terminal box. However, in the analytical calculation is assumed that the machine surface is covered entirely by semi-openfin channels. Therefore, the terminal box is neglected in the analytical calculation (Fig. 8b), and the effect of the terminal box applies as the blockage factor in the calculation process. For this case study, the machine is installed in the horizontal direction. As the axial length of the fins, spaces between the fins, and the height of the fins are different, themean value of the fins parameters of the housing is calculated. Accordingly, Table 1 shows the characteristic data of the fin section of the machine housing at a diameter of 230 (mm). To calculate the forced convection coefficient from the housing surface of the TEFC electrical machine using the empirical correlations and applying the leakage effect on the air-cooling speed during the computation process, the housing of the machine is divided into three main sections: rear section near the fan, active part, and front section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000431_ab6158-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000431_ab6158-Figure11-1.png",
        "caption": "Figure 11. Experimental test rig.",
        "texts": [],
        "surrounding_texts": [
            "In order to verify the performance of the developed control strategy in suppressing vibration, some tests have been performed using an experimental test rig like the model described in section 5. On the beam are placed two stand-alone devices (Device 1 and Device 2) and a shaker (Shaker 1), as shown in figures 11 and 12; moreover, three piezoelectric accelerometers (Acc0, Acc1 and Acc2) are mounted on the structure in order to the detect the vibration state in a global way (only for data sampling, they are not involved in the control action). The position along the beam of smart dampers, shaker and analog accelerometers is reported in table 6. As described in section 4.1, the definition of a reduced model of the system also comes from the need to limit the computational effort as much as possible. Three vibration modes have been chosen, which are modes I, II and III (17.02, 49.08 and 94.72 Hz), representing the first three resonances of the structure related to the nature of the beam. In this case, three modal coordinates represent a good trade-off between bandwidth and computational cost. A greater number of coordinates could be considered by improving the capabilities of the microcontroller. Considering the limited number of modal degrees-of-freedom, those related to inertial actuators are not considered; in this way the reduced model is guaranteed to be valid over a wider frequency range. Moreover, the contribution of information which these modes provide to the model is very low, so it is preferable to build the model with the most relevant modal components. This experimental phase, therefore, has been focused on the implementation of the Kalman filter and the Linear Quadratic Regulator on both wireless sensors, coupled with the state recovery algorithm. The experimental tests are related to the control of mode III and the disturbance is a tone whose frequency is tuned on the eigenfrequency of that mode (94.72 Hz). The results obtained in this case are shown in figure 13, in which are depicted the measured velocities related to the controlled system without (figure 13(a)) and with (figure 13(b)) the recovery of the state. The control action, in the second case, has a very high efficiency: the rms value of the velocity is in fact reduced by 94% with respect to the uncontrolled case, considering both devices involved into the control (in the first case, instead, the reduction is about 80%). In figure 14 is shown the frequency spectrum of the acceleration signals depicted in figure 13. The uncontrolled case (blue line) is related to the time domain signal up to 5 s (it is almost the same for both cases); the controlled case is divided between no state recovery (red line, figure 13(a) from 5 s on) and state recovery (green line, figure 13(b) from 5 s on). Finally, in figure 15 are represented the Frequency Response Functions of the beam to a broadband excitation generated by the shaker. The blue line shows the uncontrolled case, while the red line represents the controlled case with state recovery. The frequency range of interest varies from 20 to 280 Hz and such functions are evaluated at the Acc2 position. The most excited eigenfrequencies are related to modes III and V, due to the positions of the shaker and the accelerometer 2 with respect to the mode shapes. In the experimental case, compared to the numerical one, the covariance matrix of the error on the measurement Rv has been set up in order to obtain an estimate which is very close to the measured value; in fact, a lower error on the measurement has been considered, since the acceleration signal coming from the sensor is previously filtered. The tuning of the value of Rv has been carried out by taking one of the time histories of the velocity acquired by the sensor on the board and then by feeding the numerical model with it; at that stage, different values have been tested, until the most suitable one has been found."
        ]
    },
    {
        "image_filename": "designv11_63_0000048_j.biosystemseng.2019.10.007-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000048_j.biosystemseng.2019.10.007-Figure6-1.png",
        "caption": "Fig. 6 e MEA100T20A8B7C5D4 produced by different technologies: (a) cut by waterjet and (b) cut by laser.",
        "texts": [
            " There are some effects of the cutting technology used to produce the MEA on its behaviour in the tests: a) Laser cutting technology. High temperature is reached around the cutting area of the MEA; hence, the mechanical properties of the heated material may be altered during the production process. An annealing thermal treatment could be applied to modify the behaviour of the MEA induced by local temperature peaks due to laser cutting. b) Waterjet cutting technology. Morematerial than in case of laser cutting is removed from the MEA (see Fig. 6 for pictures of complete MEA and Fig. 7 for a detail). Hence, the rings and the arms are thinner, and the MEA requires smaller forces to deform. Additionally, it is worth to noting that the thicker MEA (4 mm) has been tempered to prevent premature fracture during the tests. A tensile testwasperformed to thedifferent selectedMEAuntil fractureswereproduced.Thetestshaveconsisted inapplyinga progressively growing tensile force in the axial direction of the MEA, while measuring both, the applied force and the axial displacement of the inner disc of the MEA with respect to the outermost ring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002739_j.matpr.2021.04.018-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002739_j.matpr.2021.04.018-Figure3-1.png",
        "caption": "Fig. 3. Deformation Plot \u2013 Isometric View.",
        "texts": [],
        "surrounding_texts": [
            "The methodology of the research work is shown in Fig. 1."
        ]
    },
    {
        "image_filename": "designv11_63_0001589_022060-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001589_022060-Figure2-1.png",
        "caption": "Figure 2: Acting forces on rotor and stator surface (light blue) and substitute forces (dark blue) [1].",
        "texts": [
            " To consider the dynamic interaction between structural elements of the generator and electromagnetic active components, a force-element has to be implemented, which can exert electromagnetic forces on structural elements of the generator. By this force-element the transient excitation of the electromagnetic field can be simulated and further studied. The force application on the generator structure needs to be implemented via a force-element in the multibody simulation environment (e.g. Simpack). Since the electromagnetic forces are a continuous stress distribution (Maxwell stresses) on the rotor, respectively stator surface, this stress distribution needs to be discretised onto subsurfaces (Figure 2, right). These discretised forces are subsequently applied onto defined points (markers) on the rotor and stator, which will distribute the force equally onto the subsurface. The electromagnetic pull between rotor and stator is decoupled by equivalent internal acting forces pushing rotor outwards and stator inwards, instead of pulling rotor and stator together, see Figure 2. [1] The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 022060 IOP Publishing doi:10.1088/1742-6596/1618/2/022060 These acting forces are influenced by the air gap width. Locations with an increased air gap width have a lower permeance and therefore experience lower field densities respectively lower forces. This leads to a significant influence of external loads, by means of eccentricities and local deformations, on the field distribution and in turn on the acting forces on rotor and stator",
            " Due to these displacement states a direct measurement of the air gap width and a direct force application between rotor and stator markers is not possible. This is due to the distance between rotor and stator markers (and the airgap width) would change with the position of the respective markers. As stated above, the angular displacement can easily be handled by decoupling rotor and stator by applying equivalent internal forces onto individual markers instead of using forces between rotor and stator markers, see Figure 2. [1] The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 022060 IOP Publishing doi:10.1088/1742-6596/1618/2/022060 The displacement states of tilting and eccentricities could be handled by calculating the air gap widths by means of cylinder equations. However, this would not consider local deformations of the rotor and stator surface, as the coordinates of the respective marker will not be in the solution room of a cylinder equation. This is exemplarily shown in Figure 4, as \ud835\udc46\ud835\udc463 is not on the same circle (cylinder surface) as \ud835\udc46\ud835\udc462"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure5-1.png",
        "caption": "Fig. 5. (a) FE reduced model [4]; (b) DoE matrix of experiments for stiffness constants calculation.",
        "texts": [
            " Anyway, an alternative strategy will also be presented in Section 4, which replaces the wire-roller-wire set by a COMBIN39 non-linear spring element (instead of MATRIX27 element), giving place to a simpler but less accurate FE modelling option. Prior to explaining all these ideas in the following sections, the calculation of the contact stiffness constants (k1,k2,k3) in Eqs. 1-6 will be briefly described now. In the previous work [4], the values of these parameters, which depend on the geometry of the contacts, were calculated from the reduced FE model shown in Fig. 5a. However, in this work, a FE analysis campaign based on the 3 level full factorial Design of Experiments (DoE) shown in Fig. 5b was performed using that FE model. As it can be seen, the variables of the DoE were the diameter of the roller Dw, the relation between wire and roller diameter \u03bb/Dw and the race factor Rf which quantifies the size of the race [12] (for large bearings, the bearing mean diameter Dpw does not affect the stiffness constants): Rf = 1+ 1 \u03bb/Dw \u2212 1 \u03bb/Dcw (7) Before the introduction of the contact stiffness results, it is worth mentioning a relevant phenomenon that takes place in the wire-ring contact. This phenomenon lies in the evolution of the wire-ring contact under the action of the load",
            " However, under the action of a high load, the gap between the wire and the ring can eventually disappear and generate the situation represented in Fig. 6b. Obviously, the closure of this gap increases the stiffness of the contacts, which can generate non-accurate results for I. Mart\u0301In et al. Tribology International 161 (2021) 107098 high loads. Fig. 7 shows the FE results and a linear approximation for the contact stiffness according to k1, k2 and k3 (see Eqs. 1\u20133), for the central case of the matrix of experiments (see Fig. 5b). Due to the non-linearity generated by the mentioned gap closure phenomenon, stiffness constants were adjusted according to the initial linear section. This means that k1, k2 and k3 are, at the very least, accurate till the 40% of the static load capacity, and conservative from then on (grey area in Fig. 7). This accuracy range is enough for the vast majority of applications considering that the design conditions are far from the static load capacity. Thus, the calculated stiffness values were fitted by means of the following expression: k [ N / m ] = c+ cDw\u22c5Dwn + c\u03bb\u22c5\u03bbn + cRf \u22c5Rf n + cDw\u03bb\u22c5Dwn\u22c5\u03bbn + cDwRf \u22c5Dwn\u22c5Rf n (8) Being: Dwn = (Dw \u2212 13)/(20 \u2212 13) (9) \u03bbn = (\u03bb/Dw \u2212 0.65)/(0.75 \u2212 0.65) (10) Rf n = (Rf \u2212 0.6)/(0.8 \u2212 0.6) (11) Where Dw and \u03bb in Eqs. 9\u201310 are in [mm], and coefficients c, cDw, c\u03bb, cRf , cDw\u03bb and cDwRf are in [N/m] and listed in Table 1. While building expression of Eq. 8, the cross term between variables Rfn and \u03bbn was found not to have significance. The relative error of this approximation was found to be lower than the 5% for any case within the defined design space (Fig. 5b). This section describes the mathematical background needed to implement the MATRIX27. As mentioned, two main mathematical operations must be carried out. For the sake of simplicity, the procedure is presented for the case of type A roller. The same steps were followed to obtain type B roller results. First, the non-linear equation system formed by Eqs. 1-6 must be transformed into a linear one in order to introduce constant coefficients into the MATRIX27 user-defined matrix. The first step consists on introducing the linear equations Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001120_j.jfranklin.2020.05.018-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001120_j.jfranklin.2020.05.018-Figure3-1.png",
        "caption": "Fig. 3. The physical meaning of protocol design.",
        "texts": [
            " ssumption 1 [17] . The sampling time T 0 is small enough in the system so that the external isturbance is a slow time-varying signal in unit time, where \u02d9 \u02dc F = \u2212 \u02d9 \u02c6 D F . (21) The parameter adaptive law of D F is designed as \u02d9 \u02c6 F = \u2212\u03b3 ( v \u2212 v d ) . (22) here \u03b3 > 0. Besides, the protocol has a clear physical meaning, and the controller of each quadrotor s constructed by the relative bearing information obtaining from the neighbors and target, so he protocol is distributed. Its physical meaning is shown in Fig. 3 . From the mathematical model of MQS, one can obtain U 1 R e 3 = m u pi + mg e 3 . Assume hat [ U x U y U z ] T = U 1 R e 3 are virtual control variables, then U x U y U z \u23a4 \u23a5 \u23a6 = U 1 R e 3 = U 1 \u00b7 \u23a1 \u23a2 \u23a3 C \u03b8C \u03d5 S \u03b8S \u03c6C \u03d5 \u2212 C \u03c6S \u03d5 C \u03c6S \u03b8C \u03d5 + S \u03c6S \u03d5 C \u03b8S \u03d5 S \u03b8S \u03c6S \u03d5 + C \u03c6C \u03d5 C \u03c6S \u03b8S \u03d5 \u2212 S \u03c6C \u03d5 \u2212S \u03b8 C \u03b8S \u03c6 C \u03c6C \u03b8 \u23a4 \u23a5 \u23a6 \u23a1 \u23a2 \u23a3 0 0 1 \u23a4 \u23a5 \u23a6 , (23) here C x and S x means function cos ( x ) and sin ( x ) respectively; U x = U 1 ( C \u03c6S \u03b8C \u03d5 + S \u03c6S \u03d5 ) ; y = U 1 ( C \u03c6S \u03b8S \u03d5 \u2212 S \u03c6C \u03d5 ) ; U z = U 1 C \u03c6C \u03b8 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001041_iet-epa.2020.0237-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001041_iet-epa.2020.0237-Figure1-1.png",
        "caption": "Fig. 1 CAD drawing of the SynRM (a) Structure, (b) Topology of the transverse-laminated SynRM",
        "texts": [
            " The passive cooling refers to the effectiveness of heat distribution within a machine, which is affected by material thermal properties, geometrical design, and interfacial thermal resistances between machine components. For example, the passive thermal design of a stator slot strongly influences the temperature gradient of the coil. Active cooling refers to the effects of heat extraction from a machine to the coolant, ultimately based on the convection phenomenon. For the research purposes, the thermal models described in this study are applied to SynRM with a four poles 10 kW, 400 V, 50 Hz, and \u2018F\u2019 insulation class. Fig. 1 illustrates the structure and topology of the machine. This machine is a TEFC machine, and it uses the air forced cooling method. Accordingly, the fin channel design is essential in the outer peripheral of the housing. An analytical model of the machine is developed, and the calculation of its components is explained. Finally, to validate the analytical model, an experimental setup and measurements are carried out. The test outcomes are used to validate the analytical data. Due to the risk of high temperature inside the slot and damaging the slot insulation, which reduces the lifetime of the insulation materials as well as the machine life, precise prediction of the temperature distribution in the machine is necessary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000041_012009-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000041_012009-Figure7-1.png",
        "caption": "Figure 7: Corona emission AC test COMSOL simulation model",
        "texts": [
            " However, the current exceeds 18 A in the second spike, which is more than the maximum current that can be drawn by the drone motor in unarmed case (i.e. without propellers). This shows the effect of the corona discharge inference on the motor speed controller performance. To validate the experiment results, a simulation model using COMSOL software was developed. A computer-aided design (CAD) model of the UAV was imported which took into consideration the different materials used in construction. The simulation model is shown in Fig. 7. The HV spherical electrode is connected to 200 kV voltage source. The distribution of the electric field over the drone surface at different azimuth angles with respect to z-axis is computed and shown in Figs. 8 - 9. In Fig. 8-a, the electric field at the UAV autopilot edges exceeds 3 MV/m for an azimuth angle = 0 degrees. In Figs. 8-b, 9-a and 9-b the electric field at the UAV motor edges is registered to be 3.4 MV/m, 5.1 MV/m and 7.5 MV/m at different azimuth angles. This shows that both the surface of the drone autopilot and motor achieve an electric field above 3 MV/m, which is the electric field magnitude required for corona onset"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000879_j.promfg.2020.04.145-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000879_j.promfg.2020.04.145-Figure2-1.png",
        "caption": "Fig. 2. Different fixation strategies of the substrate during WAAM.",
        "texts": [
            " The deposition of the beads is carried out using the element birth/death technique as also explained in detail in previous work [13]. The elements of the layers are born/activated layer by layer according to its deposition pattern. The welding direction of the alternate adjacent layer is reversed and the cooling time of about 5 s is established after each layer to improve the heat dissipation within the model. The meshing on the substrate as well as the beads is refined to enhance the precision. Five different fixation strategies of the substrate have been used in this work as shown in Fig. 2. The fixtures (clamps) are modeled in such a way that the movements of the substrate are constrained in the specific directions. The arrows on a specific fixture indicate the direction in which the substrate movement is constrained as shown in Fig. 2. The substrate is fixed under these clamps using the surface-to-surface-contact. Therefore, the substrate can bend freely under the influence of the moving heat source. In fixation strategy 1, the four edges of the substrate are clamped both in longitudinal (x) as well as in transverse (y) direction. On the other hand, in the fixation strategy 2, only the transversal edges are fixed while in strategy 3 the longitudinal edges are clamped. The purpose of using the longitudinal and transverse fixation separately in strategy 2 and 3 is to verify the side which has more influence in reducing the distortion as well as the stresses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002753_s00202-021-01301-w-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002753_s00202-021-01301-w-Figure4-1.png",
        "caption": "Fig. 4 24slots/20poles in-wheel PM BLDC motors",
        "texts": [
            " Selected in-wheel PM BLDC motor has a 3-phase trapezoidal voltage. The circular permanent magnet is segmented and placed radially and axially. The stator has double-layer winding arrangement. Table\u00a01 shows the parameters of in-wheel PM BLDC motor. Analyses were carried out on designs according to 3 different combinations of pole/slot number given in Table\u00a01. The given pole/slot number combinations are commonly used in design both in the literature and also in already manufactured motors in industry. Figure\u00a04 shows one of the designs used for the analysis. Length of magnet slot along the magnetized direction was selected 6\u00a0mm also magnet pole arc was 1620e for the magnet of this model. Also, the designs were simulated using the number of radial magnet segments from 1 to 20 in the axial direction. Each design is made using the same size of motor, material of motor, length of magnet slot along the magnetized direction, air gap length, and rated speed. (4) Wsegmented Wmonolithic = zy (E\u2215zy)2 2 ( 2a z + 2b y ) E2 2(2a+2b) = 2a + 2b z2a + y2b Eddy currents induced in electrically conducting material by a changing magnetic field, produce energy loss in the material [8, 17\u201319]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002721_tec.2021.3074818-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002721_tec.2021.3074818-Figure11-1.png",
        "caption": "Fig. 11 Prototype machines. (a) 12/10 FRPM machine. (b) Original rotor (12/10 FSPM). (c) SSRT (12/10 FSPM). (d) ASRT (12/10 FSPM). (e) EMF test. (f) Cogging torque test. (g) 3D FEA model.",
        "texts": [
            " As can be seen, even though some cases such as 9/12, 18/21 and 18/23 machines are geometrically asymmetric, the UMP is still quite low. In addition, it should be noted that Table II is also suitable for FSPM machines expect for some odd-stator slot cases since the stator slot number in FSPM machine is always even. In this section, an improved configuration for UMP reduction of FRPM machines will be proposed. As shown in Fig. 10(a), the PMs are rearranged by introducing a small space gap between the two PMs mounted on the same stator tooth. Then, the UMP can be diminished by optimizing this space gap. According to Fig. 11(a), the MMF distribution changes to: \ud835\udc39\ud835\udc5d\ud835\udc5a 2 (\ud835\udf03) = \ud835\udc39\ud835\udc5d\ud835\udc5a0 + \u2211 \ud835\udc39\ud835\udc5d\ud835\udc5a\ud835\udc5b7cos(\ud835\udc5b7\ud835\udc41\ud835\udc60\ud835\udf03) (47) \u221e \ud835\udc5b7=1 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 19,2021 at 05:25:29 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (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. where \ud835\udc39\ud835\udc5d\ud835\udc5a0 = \ud835\udc41\ud835\udc60 \ud835\udf0b \ud835\udc39\ud835\udc5d\ud835\udc5a 2 \ud835\udf03\ud835\udc5d\ud835\udc5a, \ud835\udc39\ud835\udc5d\ud835\udc5a\ud835\udc5b7 = 2\ud835\udc39\ud835\udc5d\ud835\udc5a 2 \ud835\udc5b7\ud835\udf0b [sin( \ud835\udc5b7\ud835\udc41\ud835\udc60\ud835\udefe4 2 + \ud835\udc5b7\ud835\udc41\ud835\udc60\ud835\udf03\ud835\udc5d\ud835\udc5a) \u2212 sin( \ud835\udc5b7\ud835\udc41\ud835\udc60\ud835\udefe4 2 )]",
            " According to Table II, the UMP of 6/5 FRPM machine due to PM is the main source. Therefore, it can be reduced by rearranging the PMs. Solving (49), it results \ud835\udefe4 = 10deg. In order to verify the analysis above, Fig. 10(b) compares the FEA-predicted UMP (on-load condition). As expected, the UMP either in 6/5 machine reduces sharply once the improved topology is utilized. In order to further verify the analyses above, the prototype machines are manufactured, including a 12/10 FRPM machine and a 12/10 FSPM machine with three different rotor topologies, as shown in Fig. 11. Meanwhile, the corresponding platforms are also presented in Fig. 12, which are employed to test the back-EMF and cogging torque. Fig. 11(a) shows the 12/10 FRPM machine and Figs. 11(b)-(d) are three different rotor topologies of the FSPM machine (original, SSRT and ASRT). Meanwhile, the corresponding key dimensions are listed in Table III. It should be noted these three rotors share the same stator. Besides, the back-EMF testing system is presented in Fig. 11(e), including a prototype machine, a driving motor and a platform. While the cogging torque testing system is shown in Fig. 11(f), including a prototype machine, a cogging torque meter, a driving motor and a platform. TABLE III Dimensions of the Prototype Machines Items Parameters Machine types FRPM FSPM(original) FSPM(SSRT) FSPM(ASRT) Ns/Nr 12/10 12/10 Dso/Dsi (mm) 124/73.5 128/70.4 \u03b4 (mm) 0.6 0.35 La (mm) 120 75 Br (T) 1.1 1.2 Hc (kA/m) 834 910 Shifted angle - 0\u00b0 3\u00b0 3\u00b0 By using the platform shown in Fig. 11(e), the back-EMF of 12/10 FRPM machine can be measured and plotted in Fig. 12(a). Meanwhile, the corresponding 2D and 3D superposition predicted results are also compared in this figure. The 3D FEA model is presented in Fig. 11(g). It should be noted the analyzed area is larger than the machine itself to fully consider the flux leakage around the end of PMs. Clearly, a good agreement can be obtained. Similarly, the cogging torque can be measured according to Fig. 11(f) and the corresponding results are compared in Figs. 12(b)-12(d). Both the measured and superposition predicted results indicate that these techniques can actually reduce the cogging torque. (a) (b) (c) (d) (e) Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 19,2021 at 05:25:29 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (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_63_0001656_etfa46521.2020.9211891-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001656_etfa46521.2020.9211891-Figure12-1.png",
        "caption": "Fig. 12: Comau Racer3 with defined frames",
        "texts": [
            " It is assumed that the robot\u2019s internal control is very precise, i.e. there will only be a marginal error between the real and calculated trajectory and so rE \u2248 rr applies. To calculate the force part rc of the reference signal, the differential equation KAp r\u0308c + KVp r\u0307c = \u2206F (13) is used, whereas KAp = kAp E and KVp = kVp E are the mass and damping matrices of the system. Thereby, rc is resulting from integration of (13) and so an integral term is automatically provided to control the force error [16]. Considering the robot of Fig. 12 and given reference position and orientation of the end effector, inverse kinematics is used for the calculation of desired joint positions. The inverse of the forward kinematics z\u0307E = ( IvE I\u03c9E ) = J(q) q\u0307 (14) at velocity level results in q\u0307 = J(q)\u22121(z\u0307E,d + K e) (15) using the inverse of the geometric Jacobi matrix J(q) = \u2202z\u0307E \u2202q\u0307 , desired velocities z\u0307E,d as well as position and orientation error e = (ep eo) >. Thereby, z\u0307E of (14) was replaced by (z\u0307E,d + K e) to conquer numerical drift. Using a positive definite matrix K leads to the asymptotically stable linear error system e\u0307 + K e = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001510_j.electacta.2020.136981-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001510_j.electacta.2020.136981-Figure4-1.png",
        "caption": "Fig. 4. Gradient of the magneto static field produced by the printed bonded magnet. Left column: full magnet. Right column: magnet hole-engraved at the center. Top row: at a height of 100 \u03bcm. Bottom row: at a height of 175 \u03bcm above the magnet surface.",
        "texts": [],
        "surrounding_texts": [
            "i t F p p\n0 c e t a\n2\n0 u s d\nf f c s\nm m K p a\na a p f m t u\n3\n3\nn o m t e o o t t p v i m s e\nb H t s m t\nb ( g t i o t t t m q l n a\n3 s\nfi m a i s s t t w n\nfications. Briefly, non-magnetic or magnetic screen-printed elecrodes were dip coated in 0.1 mg mL \u22121 aqueous suspension of e 3 O 4 for 5 min and after being thoroughly rinsed with DDW, they\nolarized at 1.6 V for 450 s and then at 0 V for 300 s in 0.1 M PB, H 6 containing 0.4 mM K 3 [Fe(CN) 6 ] and 0.1 M K 2 SO 4 .\nAg/Fe 3 O 4 modified SPEs were prepared either by dip coating in .15 mg mL \u22121 Ag/Fe 3 O 4 aqueous suspension for 5 min or by drop asting 10 \u03bcL 1 mg mL \u22121 Ag/Fe 3 O 4 aqueous suspension onto the lectrode surface. After drying under ambient conditions, the elecrodes were thoroughly washed with DDW. ATR-IR spectra of Fe 3 O 4 nd Ag/Fe 3 O 4 MNPs are shown in Fig. S2A.\n.6. Electrochemical measurements\nCyclic voltammograms (CVs) were recorded at a scan rate .05 V s \u22121 . Differential pulse (DP) voltammograms were recorded sing the following waveform parameters: pulse amplitude, 0.05 V; tep potential, 0.006 V; voltage step time 0.05 s. Under these conitions the effective scan rate was 0.03 V s \u22121 .\nEvaluation of the amount of Fe 3 O 4 NPs onto the electrode\u2019s surace : DP voltammograms were recorded over the potential range rom -0.4 V to 1.2 V in 0.1 M HCl. Electrodes were modified by dip oating at various concentrations of MNPs and for different immerion time intervals.\nElectro reduction of hydrogen peroxide at Prussian blue@Fe 3 O 4 odified SPEs : Cyclic voltammetry and amperometric measureents at 0 V were conducted in 0.1 M PB pH 6 containing 0.1 M 2 SO 4 in the absence or the presence of 5 mM H 2 O 2 . Calibration lots were constructed by adding 2 (5 first additions) and 5 (8 next dditions) \u03bcM H 2 O 2 in the measuring cell.\nAg/Fe 3 O 4 modified SPEs : CVs were conducted in 0.1 M PB pH 6.5 t dip coated electrodes. Riboflavin measurements were conducted t electrodes modified by drop casting in deoxygenated 0.1 M PB H 6.5 by recording DP voltammograms over the potential range rom -1 to 0 V, following a preconcentration step in an electroless\node for 1 min under stirring. The charge was calculated by inegrating the area from the onset potential in each voltammogram ntil -0.32 V with the aid of GPES software (Metrohm Autolab).\n. Results and Discussion\n.1. Magnetic characterization\nThe hysteresis loop depicts the dependence of the vector mag-\netization M (A/m) or the magnetic flux density B (T) as a function f the applied magnetic field H (A/m). A permanent magnet (hard\nagnetic material) exhibits a broad and ideally square M-H hyseresis loop. The quality of the magnet depends on the maximum nergy product (BH) max (J/m 3 ), which is represented by the area f the largest rectangle that can be fitted in the top left quadrant f the hysteresis loop [32] . The magnetic properties that can be exracted from a hysteresis loop are the saturation magnetization M S , he remanence magnetization M R and coercivity force H C . M S exresses the magnitude of the atomic magnetic moment \u03bc per unit olume. M R is the magnetization that exists at zero field and H C ndicates the resistance of a magnetized material against the de-\nagnetizing fields. In fact, M R and H C define the pattern and the ize of the hysteresis loop, which in turn, impacts the maximum nergy product (BH) max [33] .\nIn order to obtain enough magnetostatic field strength, the onded magnet disk needs to be magnetized perpendicularly. owever, its film-shape with large diameter-to-thickness aspect raio (ASR \u224855) makes perpendicular magnetization unfavorable, as trong demagnetizing fields are expected. For this shape the de-\nagnetization factor perpendicular to the disk plane is expected o be N = 0.94 [34] , which approaches the maximum value N = 1\nAs regards the fabrication parameters of the screen-printed onded magnets, the effect of NdFeB loading in the printing ink wt.%) and the number of the printing layers (L) on M R was investiated. Data in Table 1 and magnetic hysteresis loops ( Fig. 2 ) reveal hat the magnetic moment of the screen-printed bonded magnets ncreases with the thickness of the magnets [35] and the amount\nf the magnetic material in the printing ink. Data also demonstrate hat NdFeB loading plays a more pivotal role as at high loadings he magnetic flux between the magnetically isolated particles, due o the binder (graphite thermoplastic resin), is facilitated. Thicker\nagnets ( > 5 layers) and higher loadings ( > 80%) resulted in poor\nuality printings and thus were not further considered. Hysteresis\noops showing the magnetization (Am 2 /kg) of the different SPEs\normalized to the weight of NdFeB powder in the bonded magnet re shown in Fig. S3.\n.2. The shape of the printed magnet and the thickness of the ubstrate\nThe effect of the shape of the printed bonded magnet on the eld gradient as well as of the substrate thickness on the observed\nagnetic force at the electrode/electrolyte interface was evaluted by EDX spectroscopy and magnetostatic calculations. Photos n Fig. 3 A show the morphology of both the graphite electrode urface (front side) and the screen-printed bonded magnet (rear ide). The bonded magnet has a grid-like pattern which is ascribed o the weave of the polyester fabric of the screen in combinaion with the non-ideal thixotropic properties of the mixed paste,\nhich contains only 20 wt.% screen-printable ink, and to its thick-\ness. The thickness of the screen-printed bonded magnet was mea-",
            "S m F e M s n\nsured 56 \u00b1 6 \u03bcm, considerably higher compared with that of the graphite layer (11 \u00b1 1.5 \u03bcm) on the other side of the substrate.\nThe EDX mapping images of the graphite electrode surface for the non-magnetic, magnetic and hole-engraved magnetic SPEs, after they have been dip coated in 10 mL 0.1 mg mL \u22121 Fe 3 O 4 MNPs for 10 min and rinsed thoroughly with DDW are shown in Fig. 3 A. A remarkably higher amount of Fe 3 O 4 NPs is found onto magnetic\nPEs compared with that has been physically adsorbed onto non-\nagnetic SPE. For the hole-engraved magnetic SPEs, the amount of e 3 O 4 NPs is even higher and more uniformly distributed over the lectrode surface compared with the non-engraved on which the\nNPs are preferentially attracted at the periphery of the electrode urface. This is because the magnetic force on a dipole depends ot only on its magnetic moment, and the strength of the field,",
            "b b t fi c p t\no i m s c\nt s t o n t F s u m i a t m i t\n3\nF D n s d S s a c a r F s ( w s h\no\nut also on how steeply the field changes along the length of the ody, that is, the field gradient. If a body is permanently magneized, then the force will be proportional to its moment and the\neld gradient. A detailed discussion and theoretical models for the alculation of the magnetic field and of its gradient at the screenrinted (compact or hole-engraved) bonded magnet are given in he Supporting Information (paragraph S1 and Figs. S4, S5).\nFinally, photos in Fig. 3 B demonstrate the magnetic attraction f Fe 3 O 4 MNPs onto the surface of magnetic SPEs, without us-\nng an external magnetic field, after the application of 50 \u03bcL 1\ng mL \u22121 Fe 3 O 4 MNPs. For comparison purposes, a non-magnetic creen-printed electrode in which the physical adsorption of MNPs auses no discoloration of the suspension is also shown.\nSince the bonded magnet is printed at the back of the substrate, he field gradient at a height of ca. 110\u2013112 \u03bcm above the magnet urface, which corresponds to the sum of the PET substrate plus he graphite layer thickness, is of particular interest. The gradient f the magneto static field produced by the printed bonded maget at the hole-engraved and the non-engraved bonded magnet at wo different heights above the magnet surface are illustrated in ig. 4 . The 175 \u03bcm thick polyester sheet (Autostat CH7), a subtrate that we commonly use for screen printing purposes, was sed for comparison. The loci of the points where the gradient is\naximized agree with stereoscopic photos and the EDX mapping mages in Fig. 3 . These figures are produced in top-view to provide\ndirect comparison with the microscopy images and are based on he magnetostatic calculations presented in the supporting infor-\nation. However due to the cylindrical symmetry of the geometry t is more instructive to plot the fields and gradients within any of he equivalent radially directed planes, as it is done therein.\n.3. Electrochemical measurements and analytical utility\nTaken as a measure the peak current of the cathodic e(III)/Fe(II) redox transition at Fe 3 O 4 MNPs in acidic media [36] , P voltammograms illustrated in Fig. 5 manifest that at the magetic electrodes, higher loadings of Fe 3 O 4 MNPs can be achieved in horter incubation times even when SPEs are dip-coated in more iluted suspensions. In specific, when non-magnetic and magnetic PEs were dip-coated in 0.1 mg mL \u22121 Fe 3 O 4 colloidal suspenion ( Fig. 5 A), at magnetic SPEs the maximum peak current was chieved within 5 min while at non-magnetic SPEs, half of this urrent was achieved after 45 min. Comparing the peak currents t 0.2 V at the first five minutes of the dip coating process, the esponse at the magnetic SPE is ca. 4-fold higher. As shown in ig. 5 B, the magnetic SPE exhibited an almost 10-fold higher reponse when the dip coating process was performed at diluted 0.01 mg mL \u22121 Fe 3 O 4 ) colloidal suspensions, while this response\nas more than twice as high compared with the response oberved at the non-magnetic SPE that was dip-coated at a 10-fold igher concentration (0.1 mg mL \u22121 ) of Fe 3 O 4 MNPs.\nTaking advantages of the superior uptake of Fe 3 O 4 MNPs nto the electroactive surface of the magnetic SPEs, Fe O MNP-\n3 4"
        ]
    },
    {
        "image_filename": "designv11_63_0002359_jestpe.2021.3059280-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002359_jestpe.2021.3059280-Figure13-1.png",
        "caption": "Fig. 13 Schematic diagram of the special coupled inductors when the system is in (a) normal operation and (b) malfunction state.",
        "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. As analyzed above, every submotor in the multi-sector three-phase PMSM is relatively independent with little coupling inductance and isolated neutral point. Thus, one failure of submotor will not affect other submotors. Apart from the multi-sector motor, the special magnetically coupled inductors with multi-branches also provide flexible faulttolerance method. As demonstrated in the Fig. 13(a), the three-sector PMSM drive system with the proposed method is in normal operation. Since each coil is winded in the same direction, little voltage is consumed on the coupled inductors for the fundamental waves. By comparison, cut off one branch and change the carrier phase angle as / 2 when one branch A-2 malfunctions as shown in Fig. 13(b). And the phase currents in the two branches A-1 and A-3 still fulfil the basic rules mentioned above. The failure of the second branch would not interface the other two branches. Hence, the proposed system still possesses good harmonic performance as well as fault-tolerant capability. Regarding the power loss, we focus on the comparison results of topology in Fig. 1 between condition 1 (without the suppression method and coupled inductors) and condition 2 (with the harmonic suppression method and coupled inductors)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000712_auteee48671.2019.9033244-Figure20-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000712_auteee48671.2019.9033244-Figure20-1.png",
        "caption": "FIGURE 20. Temperature distribution for the stack with all three 3D Si tiers, and both diamond layers with diamond thermal conductivity set to 2000 W/m-K. The worst hot spot is 310.4 K and 7 K variation.",
        "texts": [],
        "surrounding_texts": [
            "B. CPU WITH SINGLE-TIER MEMORY The next step is to inspect the performance of the 3Dmemory integrated with the new CPU device with the hybrid silicondiamond substrate. The schematic of the structure is shown in Fig. 15. The upper and lower bounds for thermal conductivity of diamond was taken as 600 and 2000 W/m-K, respectively. The results are shown in Fig. 16 and 17. The maximum temperature on the chip is now more than 10 K lower than that obtained for 3D memory implemented with the standard CPU architecture (Fig. 10), even when the lower bound of the thermal conductivity of diamond is used in the simulation. The outline of thermal concentration of the CPU hot spots can be seen as considerably laterally dissipated in Fig. 14 as with Fig. 16 and 17, which use a diamond thermal conductivity of 600 and 2000 W/m-K, respectively. Specifically, the highest temperature in the top memory tier\nhas fallen to only 307 K with a temperature variation of only 3 K. The decrease in the maximum temperature rise is attributed to the enhanced lateral heat spreading in the diamond layer.\nC. CPU WITH DOUBLE-TIER MEMORY Finally, the last case studies the integration of two 3D memory stacks with the CPU as shown in Fig. 18. The temperature distributions of this structure are shown in Fig. 19 and 20 for diamond thermal conductivities of 600 and 2000 W/m-K, respectively. The maximum temperature increase is about 12 K.\n50 VOLUME 3, 2015",
            "Shown in Fig. 21, temperature distribution at the Cu-diamond boundary has been computed using a diamond thermal conductivity of 2000W/m-K. Results yield a temperature variance of 6 K at the diamond-Cu interface.\nTo explore this further, vertical slices were made near the hot spots of the stacked layers of two 3D tiers of memory and CPU, structurally detailed in Fig. 22. Fig. 23 and 24 illustrates the temperature distribution vertically at the same hotspot location using diamond thermal conductivities of 600 and 2000 W/m-K, respectively. Temperature distributions with both diamond thermal conductivities and without diamond are plotted vs. depth from top of the stack down, which is shown in Fig. 25. As expected, material with a lower thermal conductivity would induce a steeper slope in temperature distribution, ultimately resulting\nin a higher peak temperature. Even with diamond at its worst-case thermal conductivity, peak operating temperature still perform closer to ideal cases of diamond thermal conductivity.\nVOLUME 3, 2015 51",
            "VIII. ANALYSIS OF RESULTS Simulation results are tabulated in table 4 with and without diamond. Peak temperatures include the 30 K increase due to the HBT device. One can see that by adding a tier of memory, peak temperatures appear to increase linearly. This suggests that a CPU with two tiers of memory and without diamond is likely to have a peak temperature of 356 K (83 \u25e6C), and by aggressively thinning the Si substrate, the peak temperature can be reduced by 11 K to 345 K.\nThe most important comparison in the results is the temperature variance reduction from the use of diamond. While the variance remains roughly the same per additional stack of memory, variance has reduced by roughly 10 K with the use of diamond. With the addition of performance increase due to decreased peak temperature, reduction in temperature variance can be predicted to improve performance by asmuch as 8%.\nIX. FUTURE WORK The present results are limited by the 20 GB of memory used in the computations. It is hoped that up to 16 layers of memory over the CPU will eventually be analyzed. Further work on thermally induced stress will be examined. Additional modeling of the impact of vertical vias and details of chilled liquid cooling will be explored. Finally, revisiting the analysis in IBM\u2019s 90-nm 9HP generation will be even more attractive.\nX. CONCLUSION After previously exploring the feasibility of 3D chip stack consisting of high speed CPU and stacked DRAM, concerns were raised regarding heat removal when operating at high frequencies. Such concerns lead to processor clock frequency degradation due to reduced mobility and temperature dependent clock skew. With the use of COMSOL finite element analysis tools, diamond has proven to be potentially an effective heat spreader in a processor-memory 3D stack. Simulations have indicated an 11 K peak temperature reduction and a 10 K temperature variance reduction in the proposed CPU structure. This ultimately has led to an increase clock performance by as much as 8% due to reduced clock skew alone.\nACKNOWLEDGMENT Thanks are also given to Jerry Zimmer and Dwain Aldala at SP3, and to Professors Masashi Yamaguchi and Toh-Ming Lu at Rensselaer for extensive communications concerning Kapitza resistance mitigation. Finally, the authors would like to thank Rudolf Haring and Shurong Tian at IBM for their sight and feedback. The views presented in this paper are those of the authors.\nREFERENCES [1] G. M. Amdahl, \u2018\u2018Validity of the single processor approach to achiev-\ning large scale computing capabilities,\u2019\u2019 in Proc. AFIPS Conf., 1967, pp. 483\u2013485. [2] P. Jacob, O. Erdogan, A. Zia, P. M. Belemjian, R. P. Kraft, and J. F. McDonald, \u2018\u2018Predicting the performance of a 3D processor-memory chip stack,\u2019\u2019 IEEE Design Test Comput., vol. 22, no. 6, pp. 540\u2013547, Nov./Dec. 2005. [3] P. Jacob et al., \u2018\u2018Mitigating memory wall effects in high-clock-rate and multicore CMOS 3-D processor memory stacks,\u2019\u2019 Proc. IEEE, vol. 97, no. 1, pp. 108\u2013122, Jan. 2009.\n52 VOLUME 3, 2015"
        ]
    },
    {
        "image_filename": "designv11_63_0002356_jestpe.2021.3061663-Figure21-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002356_jestpe.2021.3061663-Figure21-1.png",
        "caption": "Fig. 21. Test rig setup. (a) Stator front view. (b) Stator side view. (c) Test bench.",
        "texts": [
            " Hence, it can be found that the back-EMF amplitude of CPM-II-B (two staggered segments) is larger than that of the CPM-II-C (three staggered segments), and CPM-II-A has the lowest back-EMF amplitude. Additionally, the corresponding torque performance is compared in Fig. 20 (b). It is noticed that the CPM-II has the highest torque capability. However, the average torque of CPM-II-B is larger than that of CPM-II-A and CPM-II-C. The FEA results are consistent with the above theoretical analysis. Based on Tables I and II, a prototype is manufactured to validate the aforementioned analysis, as shown in Fig. 21 (a) and (b). The test bench was built, as shown in Fig. 21 (c). Fig. 22 shows the waveforms of phase back-EMFs under the rated speed. It can be noticed that the phase back-EMFs present good symmetry. Furthermore, the measured result is compared with the FE predictions in Fig. 23. The amplitudes of the measured and FE-predicted fundamental back-EMFs are 44.81 V and 45.53 V, respectively. The minor error (1.6%) is within the acceptable range. The amplitude of the 3th harmonic of the measured and FE-predicted results are 1.23 V and 1.21 V, respectively. The back-EMFs at different rotating speeds were measured and compared in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000795_ec-03-2019-0083-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000795_ec-03-2019-0083-Figure2-1.png",
        "caption": "Figure 2. The simplified model of gear tooth using Ishikawamethod",
        "texts": [
            " In this study, the meshing stiffness of the healthy gear pair is analyzed through the Ishikawa method, which is verified with the energy method. 3.1.1 Calculation meshing stiffness with Ishikawa method. According to Ishikawa method, gear tooth is usually simplified as a cantilever beam that is combined with a Locomotive traction gear pair system trapezoid and a rectangular, in which the length of rectangular is defined as the dangerous section of gear tooth that is calculated by 30-degree tangent method, as shown in Figure 2. The deformation of the single tooth along the direction of the meshing line can be expressed as: d \u00bc d Br \u00fe d Bt \u00fe d S \u00fe d G (1) where d Br, d Bt, d S and d G represent the bending deformation of the rectangular portion, bending deformation of trapezoidal part, deformation produced by shear and the deformation caused by the sloping of the basic part, respectively, which can be represented by: d Br \u00bc 12Fncos2v x EBs3F hxhr hx hr\u00f0 \u00de \u00fe h3r 3 d Bt \u00bc 6Fncos2v x EBs3F hi hx hi hr 4 hi hx hi hr 2ln hi hx hi hr 3 hi hr\u00f0 \u00de3 d s \u00bc 2 1\u00fe \u00f0 \u00deFncos2v x EBsF hr \u00fe hi hr\u00f0 \u00deln hi hr hi hx d G \u00bc 24Fnhx cos2v 2 pEBs2F (2) Here E, B and v refer to Young\u2019s modulus, face width and Poisson\u2019s ratio, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure5-1.png",
        "caption": "Fig. 5. The shape plot of temperature distribution at the starting point of laser beam.",
        "texts": [],
        "surrounding_texts": [
            "From the experimental results and the discussions, the following conclusions can be drawn: i. The deposited clad of the composites varied with the change in the selected process parameters, this in turn affected the physical, mechanical, and metallurgical properties of the composites. ii. From the analysis of the results obtained from characterization, the combination of laser power of 900 and 1000 W, powder feed rate of 2.5 g/min, gas feed rate of 2.0 L/min, spot diameter of 2 mm and scanning speed of 0.8 and 1.0 m/ min produced composites with an increased percentage of dilution rate, aspect ratio and enhanced powder efficiency. Based on the analysis of these findings, these process parameters are optimum for producing highly reinforced composites. iii. The geometrical characteristics of the composites were influenced by the laser power, scan velocity and the powder feed rate. Finally, from the computational modelling, the movement of laser from the starting point to the end changes the temperature distribution and enhanced microstructures in the melt pool are depended on the laser input, scan velocity and the faster cooling rate. CRediT authorship contribution statement Dr. Olawale S. Fatoba is the co-supervisor of the Master student that work on this work and he played a crucial role in the modelling and thorough supervision of the work. Mr. Adedoyin M. Lasisi is the Master\u2019s student that worked on this project. He carried out characterizations and experimental procedures of the work. He writes the literature and did carry out the validation of the work. Dr. Omolayo M. Ikumapayi is a project administration, he helps in the drafting of the literature and in put the entire work in a right journal formal and revising the work. He revised and did the editing of the work. Dr. Stephen A. Akinlabi helps in proof-reading the entire manuscript and also helps in reviewing the work and contributes immensely in the methodology and discussions of the results. He also provides additive materials. Prof. Esther T. Akinlabi happens to be the main supervisor of the Master\u2019s student that did this work. She provides funding for materials and software and She conceptualized the project and did supervision for the success of the research. 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."
        ]
    },
    {
        "image_filename": "designv11_63_0003200_s00419-021-01997-z-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003200_s00419-021-01997-z-Figure2-1.png",
        "caption": "Fig. 2 Scissor deployable structure composed of N \u00d7N SLEs in the ANCF frame",
        "texts": [
            " In addition, the cross-section of beam element in the proposed model is not only represented by gradient r,y and coordinate y, but also by higher-order terms of coordinate x and y. Furthermore, the gradient vector r,x is the first-order term of the transverse coordinate x and the cubic term of the radial coordinate y, and the gradient vector r,y is the square term of the coordinate y. Therefore, the cross-section of the beam element becomes a curved surface, which can produce warping deformation. 3.1 Modeling of Scissor-like element The basic scissor element is shown in the upper left part of Fig. 2. Each basic unit consists of two bars, which are hinged at the midpoint by a pin. Different SLEs are connected by pins at the end nodes. According to the assembly direction and angle, various kinds of deployable structures that can realize different functions can be established, such as linear deployable structures, square deployable structures, butterfly deployable structures. In Fig. 2, each bar can be considered as a beam with the same geometric and material properties. Based on the absolute coordinate method and the improved beam element model, the position of any point of bar AB in the beam element is expressed as, r Se S [ rA rAy ro roy rB rBy ] (11) where S is the shape function of the beam element, and e is the node coordinate vector of the element. ri (i A, o, B) is the global position coordinate of the node, and riy is the gradient of the coordinate relative to the Y-direction",
            " These unfavorable factors not only reduce the kinematic stability of the mechanism, but also brings great difficulties to the dynamic analysis and solution process. In view of this, according to the configuration characteristics of the deployable structurewith SLEs, a node separationmethod is adopted by reassembling the position coordinates and gradient vectors inside SLEs and between the adjacent elements,which can eliminate the constraint equations and reduce the difficulty of solving of multibody system. Take a planar array deployable structure (N \u00d7 N ) as an example, as shown in the right half of Fig. 2. In a SLE, the coordinates of two bars assembled in the absolute node coordinate frame are given by qs [ rA rA,y ro ro,y rB rB,y rC rC,y ro\u2032 ro\u2032,y rD rD,y ] (21) where ri ( i A, B,C, D, o, o\u2032) is the position vector of the i node, and ri,y is the gradient vector of the i th node relative to y-axis. Using the transformation matrix T1, the coordinates qs can be transfer as: qs1 T1qs [ rA rA,y rB rB,y ro ro,y ro\u2032 ro\u2032,y rC rC,y rD rD,y ] (22) In which, T1 can be obtained by filtering method. Suppose it is a n \u00d7n all-1 matrix, and the elements in the matrix are determined by the following constraint relations",
            " 7 that the warpage of the section increases with the increase in the load, which proves the effectiveness of the proposed beam element in calculating the warpage of the section. At the same time, it can be predicted that the greater the flexibility of the structure, the greater the cross-section warpage, and the two should be in a positive proportion. Therefore, for the large deployable structure, it is necessary to use the improved ANCF beam element to calculate its dynamics because of its large member size and high flexibility ratio. 4.2 Structural dynamics with planar deployable structure According to the array direction shown in Fig. 2, a 2\u00d72 planar deployable structure composed of 4 SLEs is selected as an example to analyze its modal change. Each bar has the same length, cross-sectional area and material. The length of the bar is 50 mm, Young\u2019s modulus E 70\u00d7 109 N/m2, density is 2.77\u00d7 103 Kg/m3, and moment of inertia is 1.66 \u00d7 10\u221211m4. In addition, the height and width of cross-section are 0.003m and 0.004m, respectively. When the initial deployment angle is 60\u25e6, the natural frequency obtained by MATLAB programming is compared with those obtained by ANSYS simulation, as shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002423_012128-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002423_012128-Figure4-1.png",
        "caption": "Figure 4. Assembled re-entrant cubic unit cell. a) Front view, b) Isometric view.",
        "texts": [
            " The cubic unit cell (Figure 1) that resembles a systematically collapsed cube is re-entrant in nature. This cell can be divided into repeating units that can be further subdivided into standard geometric shapes. Based on the availability of standard sized raw materials aluminium cubes (6.1 mm) and steel links (\u00d8 1.3 mm) are selected. These are fixed parameters in the unit cell. The repeating unit (Figure 2) consists of 5 steel links assembled in a particular formation to facilitate assembly with other repeating units. The re-entrant cubic unit is an assembly of 6 repeating units (Figure 4). The dimensions of the unit cell assembly can be controlled with two parameters, length of the steel link (\ud835\udc59) and angle (\ud835\udf03) between the steel links (Figure 2). ICRIET 2020 IOP Conf. Series: Materials Science and Engineering 1070 (2021) 012128 IOP Publishing doi:10.1088/1757-899X/1070/1/012128 These parameters are assigned 3 levels to determine the optimum combination (Table 1). The desired functionality of the unit cell is to provide the maximum negative Poisson\u2019s ratio with minimal induced stress, while subjected to an external force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002806_tia.2021.3079169-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002806_tia.2021.3079169-Figure1-1.png",
        "caption": "Fig. 1. Proposed single-drive bearingless motor for cooling fan applications.",
        "texts": [
            " Consequently, the proposed method of passing through the critical speed by high acceleration has been clarified. Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 16,2021 at 06:27:18 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (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. II. PROPOSED MACHINE STRUCTURE Figure 1 shows a cut view of the fabricated prototype machine of the proposed single-drive bearingless motor by the authors for cooling fan applications [18]. This machine has a single-drive bearingless motor and two sets of repulsive passive magnetic bearings. At the top of the rotor shaft, a fan blade is installed. At the bottom, a displacement sensor is installed to detect the rotor axial displacement. One set of three-phase winding is wound in the center stator core as the concentrated winding. Figure 2 shows the xy cross-sectional view of the singledrive bearingless motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000116_012069-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000116_012069-Figure5-1.png",
        "caption": "Figure 5. Sliding Surface or Manifold of the Sliding Mode Controller",
        "texts": [
            " PIV controller is represented by Equation (7), Ns Ns KsE s KsEKsU vp i )()()( (7) where , KP - Proportional Gain constant Ki - Integral Gain constant Kv-Velocity constant N-constant of the feed forward control E(s) - Error in s domain Like the PIV controller, there also exist three parameters for sliding mode control. The parameters are sliding mode gain, K, sliding mode slope, \u03bb and sliding mode thickness, \u03d5. These parameters are used to control the system according to the sliding manifold or surface shown in Figure 5. As the sliding mode thickness, \u03d5 increases, the chattering of the system response decreases. However, the steady state error slightly increases as the thickness, \u03d5increases. Equation (8) gives the function of the sliding surface, ees (8) The sliding mode function, s is the sum of the products of slope, \u03bb and error, e and derivative of error. The derivative of the error is needed to do the tracking of a set point. Equation (9) shows the switching function of the sliding mode controller, sfor s sforssign FunctionSwitching )( (9) When |s| is greater than\u03d5, the signum function of s is considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000284_j.compeleceng.2019.106534-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000284_j.compeleceng.2019.106534-Figure1-1.png",
        "caption": "Fig. 1. Motor model structure.",
        "texts": [
            " In Section 2 , the BIM model and field-circuit coupling theory are introduced. In Section 3 , the motor starting characteristic is analyzed and demonstrated by simulation in ANSYS Maxwell . In Section 4 , its load characteristic is similarly analyzed and simulated. In Section 5 , the suspension performance is investigated. In Section 6 , two prototypes are built, and their speed response and radial displacement are tested. In Section 7 , the conclusions based on the analytical method, simulations as well as experiments are drawn. Fig. 1 shows a two-dimensional model of the motor built by ANSYS Maxwell . The stator adopts double-layer windings and the slot type is pyriform. The rotor has cast aluminum and the slot type is parallel. The stator and rotor core are laminated by silicon steel sheets, with a lamination coefficient of 0.98. Table 1 lists the parameters of the BIM. Fig. 2 illustrates the distribution of torque winding and suspension winding in the stator. The pole-pair number of torque winding is 2 and the phase sequence is + A 1 , \u2212C 1 , + B 1 , \u2212A 1 , + C 1 , \u2212B 1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003002_s10846-021-01416-z-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003002_s10846-021-01416-z-Figure4-1.png",
        "caption": "Fig. 4 Link and joint parameters with respect to the defined frame.",
        "texts": [
            " At this point, if the \u0394t is small enough, and H (t) does not vary largely during the\u0394t times, then H\u0302 (t) can be rewritten: H\u0302 (t) \u2248 H (t \u2212 \u0394t) = x\u0307(t \u2212 \u0394t) \u2212 B\u0304u(t \u2212 \u0394t) (10) Therefore, by combining (10) with (9), the time delay control law can be obtained as follows: u(t) = u(t \u2212\u0394t)+ B\u0304 + (x\u0307d(t)\u2212 x\u0307(t \u2212\u0394t)\u2212Ame(t)) (11) If \u2225\u2225\u2225I \u2212 BB\u0304 +\u2225\u2225\u2225 < 1, the system is stable. More details can be found in [29]. A commercial excavator can be considered that is a serial robot with three degrees of freedom and has unknown and unpredictable dynamic, friction, and disturbance modeling et al. It is meaningless to take into account the moment of inertia and center of mass since electric pressure controller and pressure sensor are not used. Only kinematic information is given. Figure 4 shows the frame, link, and joint parameters. Based on unknown parameters, the general dynamic modeling derived by Lagrange formulation is as follows: M(q)q\u0308 + C(q, q\u0307)q\u0307 + g(q) + d + f = u (12) where M(q) \u2208 R3\u00d73 is the unknown moment of inertia matrix, C(q, q\u0307)q\u0307 \u2208 R3\u00d71 is unknown centrifugal and Coriolis terms, and g(q) \u2208 R3\u00d71 is unknown vector of gravity forces; d and f are unknown or unpredictable disturbance and friction vectors respectively, u \u2208 R3\u00d71 is an external input vector, and qT = (q1, q2, q3) T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001497_kem.861.113-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001497_kem.861.113-Figure7-1.png",
        "caption": "Fig. 7 Finite element simulation of turbine blade under thermo mechanical fatigue load",
        "texts": [
            " [16] conducted thermomechanical fatigue experiments under different stress/temperature range, relaxation time and phase angle, studied the thermomechanical fatigue behavior, failure mechanism and life prediction method of the alloy. Through the comparative analysis of the fracture, it was found that the same phase thermomechanical fatigue mode was creep fatigue failure, and the different phase thermomechanical fatigue mode was oxidation fatigue failure. In addition, they also established a life prediction model based on critical plane, which has strong universality, simple mathematical form, good agreement between prediction results and test data, and has a certain practical value (Figure 7). Liang et al. [17] of Tsinghua University observed and measured the growth behavior of small cracks in nickel base single crystal superalloy from room temperature to 980\u2103, analyzed its micro failure mechanism, determined the nucleation position, growth rate and growth behavior of cracks, and the interaction between cracks and inclusions, gas holes or grain boundaries. It was found that the traditional Paris law can no longer describe the propagation behavior of small cracks. The service environment of aeroengine turbine blade is very bad"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002987_s40964-021-00199-x-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002987_s40964-021-00199-x-Figure13-1.png",
        "caption": "Fig. 13 Preliminary segmentation of the benchmark with l = 15\u25e6",
        "texts": [
            " The prediction is also in this case supported by the observation of the manufactured part shown in Fig.\u00a012b. The manufactured part exhibited capillarity in the upper and rear parts, as shown in Fig.\u00a012. The calculated value of da in these regions was equal to 1 (see also Fig.\u00a011c, which confirms the adequacy of this predictor). Finally, Fig.\u00a011d and e shows the calculated maps of dsk and dth , respectively. No critical regions are observed for these defects. Accordingly, no surface sinking and thermal bleeding were found in the manufactured part. Figure\u00a013 shows the clusters of elements obtained using the procedure described in Sect.\u00a03.1 with a limit angle l = 15\u25e6 . The seed positions have been assigned randomly. The clusters obtained with a different choice of seed position do not show significant differences from the ones shown in Fig.\u00a013. The benchmark part does not have specific functional requirements, since it has only the scope of emphasising the process-induced defects. Therefore, the objectives of the optimisation aim at only eliminating the two main defects highlighted in the previous section. For the scope, two clusters of mesh elements are defined, as shown in Fig.\u00a014. Cluster 1 (Fig.\u00a014a) comprises 6656 elements including the roof, the front deck, and the rails. The relative importance of stair-casing wst is equal to 10 within this cluster, whilst the weights of the remaining defects are equal to 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure25-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure25-1.png",
        "caption": "Fig. 25 Stress analysis of the tower",
        "texts": [
            " The EBDOP-1 and EBDOP-2 designs both show an ~ 336\u00a0MPa decrease at the outermost attachment points. This is the result of adopting a single pressing, which provides better local stiffness. However, additional hotpots above 350\u00a0MPa that do not appear on the baseline are observed. The software-based optimised design SBDOP shows an ~ 40\u00a0MPa increase in stress at the right-hand 1 3 side tower. The EBDOP-1 and EBDOP-2 designs show an ~ 314\u00a0MPa decrease at the outermost attachment points. This is again the result of single pressing. The stress analysis of the tower is shown in Fig.\u00a025 and presented in Table\u00a07. Expert-based optimisation-1 (EBOP-1) is better method to reduce stress for tower. It is shown that the Software-based optimised design SBDOP results is better for mass reduction. Final results obtained from topology optimisation are shown in Table\u00a08. By considering Software-based optimized results, total mass reduced from 12.92 to 11.42\u00a0kg. Additionally, lower panel thickness reduced from 1.5 to 1.0\u00a0mm. Softwarebased optimisation method (SBOP) is better method to reduce total mass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002963_16878140211028448-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002963_16878140211028448-Figure10-1.png",
        "caption": "Figure 10. The mechanical structure drawing of the simple jump experimental platform.",
        "texts": [
            " Owing to this, the robot takes a longer time to take off, which is confirmed by the Figure 6. On the other hand, the maximum vertical velocity of the CoM is also smaller, which is confirmed by the Figure 7. For the same reason, the joint angle of precise model changes less than the ideal model, which is confirmed by the Figure 8. The experiment of the vertical jump is carried out on a simple jump experimental platform, which was built in terms of BHR-6 simplified structure. The mechanical structure drawing of it is shown in Figure 10. Parker motors K044100-6Y with reduction ratio of 100 is installed in each joint. To enhance the structural strength of the robot, structural reinforcement parts are added between the left and right leg plates. Meanwhile, the arc design on the leg board improves the motion range of joints. The parameter of the simple humanoid robot jump experimental platform is shown in Table 3. The control system needs to coordinate multiple motors to work together. At the same time, in order to ensure that the trajectory executed by the joint meets the desired dynamic characteristics, the control period needs to be as short as possible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure12-1.png",
        "caption": "Fig. 12. Half bearing simplified model (2 M DoF).",
        "texts": [
            " In this last aspect, only one sector with two roller halves was modelled for the axial load case, taking advantage of cyclic symmetry; for radial and tilting moment load cases, only half bearing was modelled, since the geometry, the boundary conditions and the loads are symmetric. The geometry of the studied bearings is shown in Fig. 11, with the geometrical parameters of Table 2 [4]. Rings were dimensioned following the standard geometry proposed in [18], but dependent on the housing H rather than on the rolling element diameter Dw [12]. Fig. 12 shows the efficient model developed in this work. The diagonal lines represent the MATRIX27 or COMBIN39 elements, which are connected to the rigid shells in the contact surfaces (in red) by means of rigid beam elements. It can be observed that these models do not require refined meshes which, together with the fact that rollers and wires are not considered, significantly reduces the DoF and the analysis cost. However, this does not mean that mesh dependency have no influence. Mesh elements should have good aspect ratio and an adequate mesh size in the contact zone that allow defining a restrictive contact penetration tolerance",
            " The final mesh was the result of a mesh sensitivity study, where it was proven that this strategy retrieved the same results with fewer elements than the equivalent model performed with a continuous mesh; element sizes are also optimal, since smaller elements did not retrieve more accurate results. Contacts are grease lubricated, so they were defined as frictional with a friction coefficient of 0.1 [19\u201321], updating the contact stiffness each iteration to achieve 1 [\u03bcm] of maximum mesh penetration. As a result, the half-bearing reference FE model has 26,914,527 DoF and a large number of non-linear contacts, while the efficient model shown in Fig. 12 only has 2,054,6252054,625 DoF and no non-linear contacts defined. For the sake of consistency, the reference model uses the same meshing strategy than the one in the previous work [4]. Load, symmetry and boundary conditions in the reference FE model were the same as in the efficient model, shown in Fig. 12. Next, the results of the validation process are presented. As mentioned, pure axial, radial and tilting moment load cases were studied, and overall stiffness and local contact results were reported as indicators of the structural performance of the bearing. In order to assess the computational cost of the different models, Table 3 gathers the CP and elapsed time of each simulation. All the I. Mart\u0301In et al. Tribology International 161 (2021) 107098 analyses were performed in a workstation with an Intel\u00ae Xeon\u00ae E5\u20132697 v3 @ 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002315_j.matpr.2020.12.487-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002315_j.matpr.2020.12.487-Figure4-1.png",
        "caption": "Fig. 4. (a) Design of foot insole; (b) render image of foot insole.",
        "texts": [
            " From the calcaneus inferior base to the central Axis of tibia the angle is 74 this indicates flat foot [5] as shown in Fig. 3(B). Foot is one of the complex structure in human anatomy. Skin data is taken for reverse engineering. Foot length of 244.47 mm represents insole size as 7 according to English system [6]. CATIA software is used to design foot insole for case II considering multiple curve sections of foot to develop top face of insole. The front face is designed according to rocker curve structure, stiffener at arch and minimal curve at rear as shown in Fig. 4. Analysis is performed with different weight bearing conditions of foot on designed insole, modified insole which is initially optimized with solid data. Further optimization is performed analyzing with different lattice structures. Lower limb dimensional relationships are key for identifying the body height based on foot length. Parameters belonging to the right foot of a female are considered in this work as the value of constant, regression coefficient varies with gender and foot. d 2. (b) subtalar joint angle for case 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001662_012010-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001662_012010-Figure1-1.png",
        "caption": "Figure 1. Electrical transmission inspection robot scheme [11].",
        "texts": [
            " Inspection and maintenance of electrical transmission lines are necessary and done regularly according to the standards of each country [1\u20132], since any problem during the transmission of the electrical energy lines to cities, industries, etc. may affect human life. Usually, inspection and maintenance of electrical transmission lines are done by human forces. By this, human forces are located in a potential hazard environment. Therefore, to reduce this risk, different types of robot inspectors are designed and used in some regions [1\u201310]. Each of these robots is capable of doing different tasks. One of these robots shown in figure 1 [11]. Executive and mathematical modeling experience shows that even with a constant speed movement of the robot over the line, vibrations occur [3\u201310, 12\u201316]. These vibrations can cause damage as well to the electrical line as to the robot. Therefore, it is needed to conduct mathematical modeling for different situations that may robot faces [11\u201316]. This article, which is the extended work of [12\u201315], deals with the construction of mathematical models for a situation in which the electrical line stretched and supported by two isolators, one horizontal and the other vertical and robot inspector moves with a constant speed on it"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000795_ec-03-2019-0083-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000795_ec-03-2019-0083-Figure1-1.png",
        "caption": "Figure 1. Structural schematic (a) and the corresponding dynamic model (b) of locomotive traction gear transmission system",
        "texts": [
            " Subsequently, the dynamic equations of the locomotive traction gear system are derived by Newton\u2019s laws of motion in Section 5. Section 6 discusses the non-linear dynamic responses for the traction system of the locomotive with the influences of control parameters. Finally, some brief conclusions are presented in Section 7. The traction gear system of a representative locomotive mostly consists of traction motor, coupling, driving gear, driven gear, wheelset and so forth, which can be seen in Figure 1(a). EC As the power part of locomotive, the traction system is used to transmit the driving torque generated by traction motor to wheelset via gear-pair, which makes the locomotive running or braking. To research the dynamic behavior of the traction gear system conveniently, some assumptions are provided as follows: The meshing gears are involute spur gears, in which meshing force is simplified as the spring-damping force and is always in the meshing plane; The transmission shafts and bearings that support gears are inflexible; and The gears are assumed to be rigid, where only the torsional displacement of gear system is considered. Then, the simplified traction system model is illustrated in Figure 1(b). In the dynamic model, u i, Ii and rbi (i = 1, 2) are the angular displacements, the moment of inertia and radius of base circles of the driving and driven gears. cm and k(t) are the meshing damping and stiffness of gear-pair, respectively. Ff stands for the sliding friction force between the two gears. In addition, T1 is the driving torque of the gear system, which is considered to keep constant. T2 represents the load torque, namely, adhesion torque caused by adhesion characteristics between wheelset and rail",
            " Sketch of contact condition between wheel and rail including pure adhesion zone (a), coexistence zone of adhesion (b) and slip (c), and pure slip zone (d) Slip zone Adhesion zone OA AB BC (a) (b) (c)(d) Figure 10. Adhesion coefficient curve between wheel and rail A d h es io n c o ef fi ci en t \u03bc Creep ratio s Actual application curve Kalker theory curve O A B C D EC Fw \u00bc mQ T2 \u00bc FwR (27) WhereQ represents the axle load of the wheelset. Substituting equation (26) into equation (27), the load torque T2 can be transformed into the following expressions: T2\u00bc mmRQs=sm Case I mmRQ\u00fe RQkm s sm\u00f0 \u00de Case II ( (28) Figure 1 presents the simplified torsional lumped parameter model of the traction system. Thus, the differential equations of motion of the system can be described as follows: I1 d2u 1 dt2 \u00bc T1 rb1 XN j\u00bc1 Fj Tf1 I2 d2u 2 dt2 \u00bc T2 \u00fe rb2 XN j\u00bc1 Fj \u00fe Tf2 8>>>< >>>: (29) The relative displacement x along the meshing line with x = rb1u 1 rb2u 2 e(t) is introduced, so that the equation (29) can then be expressed as: d2x dt2 \u00fe 1 me1 \u00fe 1 me2 XN j\u00bc1 Fj \u00fe 1 rb1me1 Tf1 \u00fe 1 rb2me2 Tf2 \u00bc 1 rb1me1 T1 \u00fe 1 rb2me2 T2 d2e t\u00f0 \u00de dt2 (30) Whereme1 \u00bc I1=r2b1 andme2 \u00bc I2=r2b2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure7-1.png",
        "caption": "Figure 7. Stress-strain state at the predicted moment of shank tear-off\u0430",
        "texts": [
            " 6 the bushings are displaced (second and third position). This can be explained by the distortion of the plane of contact of the package with the upper end of the bushing as a result of the sheet tensioning, as well as by the fact that the bushing is planted on a rod with a gap. The displacement of the bushing and its inclination favors the growth of the tension due to the weakening of the support. Further retracting the sleeve position is slightly aligned, but the wedge-shaped gap between the sleeve and the package does not disappear. On the fig. 7 the field of equivalent voltages is shown at the moment when the shank of the rod is to be detached. This torque is determined by the value of the stresses in the stem neck at the level of 1200 MPa, which corresponds to the strength limit according to the normative and technical documentation on the bolt rivet rods made of BT16 material. Taking into account the neck diameter of 1.8 mm and the neck area of 2.54 mm2, the tear-off force can be approximately determined as 2.54*1200 \u2248 3000 N. Therefore, it can be recommended that the axial force on the tool cam should not exceed 3000 N. At the same time, if the axial force is less than 3000 N, the desired axial interference will not be provided, as can be seen from the figure. 7 and 8. On the fig. 7 it is visible that tensions in a rod in a contact zone with a package make from 300 MPa on edge of a head to 1000 MPa on a site of transition of a head to ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 a cylindrical part. In the sheet there are stresses of about 200 MPa. In the previous moment of calculation, differing on axial position of a cam on 0,11 mm, tensions in a rod in a contact zone with a package make from 100 MPa to 800 MPa, that is appreciably less, as shown on fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002065_j.matpr.2020.11.262-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002065_j.matpr.2020.11.262-Figure7-1.png",
        "caption": "Fig. 7. The shape plot of temperature distribution towards the end of the laser beam movement.",
        "texts": [],
        "surrounding_texts": [
            "From the experimental results and the discussions, the following conclusions can be drawn: i. The deposited clad of the composites varied with the change in the selected process parameters, this in turn affected the physical, mechanical, and metallurgical properties of the composites. ii. From the analysis of the results obtained from characterization, the combination of laser power of 900 and 1000 W, powder feed rate of 2.5 g/min, gas feed rate of 2.0 L/min, spot diameter of 2 mm and scanning speed of 0.8 and 1.0 m/ min produced composites with an increased percentage of dilution rate, aspect ratio and enhanced powder efficiency. Based on the analysis of these findings, these process parameters are optimum for producing highly reinforced composites. iii. The geometrical characteristics of the composites were influenced by the laser power, scan velocity and the powder feed rate. Finally, from the computational modelling, the movement of laser from the starting point to the end changes the temperature distribution and enhanced microstructures in the melt pool are depended on the laser input, scan velocity and the faster cooling rate. CRediT authorship contribution statement Dr. Olawale S. Fatoba is the co-supervisor of the Master student that work on this work and he played a crucial role in the modelling and thorough supervision of the work. Mr. Adedoyin M. Lasisi is the Master\u2019s student that worked on this project. He carried out characterizations and experimental procedures of the work. He writes the literature and did carry out the validation of the work. Dr. Omolayo M. Ikumapayi is a project administration, he helps in the drafting of the literature and in put the entire work in a right journal formal and revising the work. He revised and did the editing of the work. Dr. Stephen A. Akinlabi helps in proof-reading the entire manuscript and also helps in reviewing the work and contributes immensely in the methodology and discussions of the results. He also provides additive materials. Prof. Esther T. Akinlabi happens to be the main supervisor of the Master\u2019s student that did this work. She provides funding for materials and software and She conceptualized the project and did supervision for the success of the research. 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."
        ]
    },
    {
        "image_filename": "designv11_63_0000068_edpe.2019.8883869-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000068_edpe.2019.8883869-Figure6-1.png",
        "caption": "FIGURE 6. Winding arrangement of the motor prototype.",
        "texts": [
            " Considering all the analysis above, the 48-slot 44-pole motor is the best scheme, and its vibration is the most satisfying. 8540 VOLUME 4, 2016 V. PROTOTYPE DESIGN AND EXPERIMENT A. PROTOTYPE DESIGN AND MANUFACTURE The scheme of 48-slot 44-pole motor is selected as the best one, and its final parameters are listed in Table IV. Open slot and double-layer winding is adopted in the stator, and the permanent magnet material is N35UH. The type of silicon steel sheets used in both stator and rotor is 50DW310-Z. Ansoft software is used to simulate the motor. The winding arrangement is shown in Fig.6. Prototype is manufactured according to the results of electromagnetic calculation, and the motor structure is shown in Fig.7. There are 5 permanent magnets on rotor in axial direction of each pole and each permanent magnet is divided into 10 small permanent magnets in order to decrease the influence of eddy current. B. EXPERIMENT AND ANALYSIS While doing the experiments, the motor is supplied by a six-phase inverter, and the load of motor is a six-phase PM motor. The test system is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002323_978-3-030-67411-3_32-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002323_978-3-030-67411-3_32-Figure13-1.png",
        "caption": "Fig. 13. Group C\u2019s Drawing Robot (mini-project 1) and Dancing Humanoid (final project) Fig. 14. Group C\u2019s Drawing Robot (mini-project 1) and",
        "texts": [],
        "surrounding_texts": [
            "The flow state is defined byCs\u00edkszentmih\u00e1lyi [15] as the state that people are so absorbed in a challenge with which the optimal experience itself becomes enjoyable enough to make them feel that the experience itself gives them enough reward to continue the activity. It is indicated as one of the important concepts in gamedevelopment as explained by Prensky [16]: Achieving a true \u201cflow\u201d experience, however, requires a second factor besides just \u201cleveling up,\u201d a factor that complex games also provide. That factor is always being kept in a narrow zone between the game\u2019s being too hard (\u201cI give up\u201d) and too easy (\u201cI\u2019m not challenged at all.\u201d) (p. 59). To keep the flow state, the skill level required to successfully solve a presented challenge needs to be at a manageable level, not too hard or too easy [17]. The flow state creates an incentive for a person to learn new skills, which contributes to his/her engagement in increasingly difficult challenges.Cs\u00edkszentmih\u00e1lyi explains thatwe are all capable of reaching the state of flow. It also provides mental energy promoting attention"
        ]
    },
    {
        "image_filename": "designv11_63_0000517_3352593.3352679-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000517_3352593.3352679-Figure2-1.png",
        "caption": "Figure 2: Geometry of 3-DOF planar cable-driven parallel robot",
        "texts": [
            " To obtain the workspace, generally branch-and-prune algorithm is used [20], which essentially employs the test to detect if the box lies fully outside the WFW (using consistency techniques), the test to detect if the box lies fully inside the WFW (using feasibility algorithm) and the bisection algorithm. Since the key interest of this paper is to demonstrate the application of improved closed form algorithm instead of simplex algorithm as feasibility algorithm, the workspace, in this paper, is simply obtained by dividing the searchspace into grid of boxes and then testing each box if it is fully inside the WFW using feasibility algorithm. Consider a planar 3-DOF cable-driven parallel robot driven by 4 cables in crossed configuration as shown in Fig. 2. The base frame and the mobile platform of the cable robot are a square of side 1 m and 0.2 m respectively. The orientation of the mobile platform is defined by angle \ud835\udf03 and is taken positive in the anti-clockwise direction. The tension limits in each cable for WFW computation are defined by interval [1, 50] while the force acting on mobile platform is defined by interval vector [f] = [[-10, 10], [-10, 10], [- 0.5, -0.5]] (N, N, N.m.). The constant orientation workspace for 2 different cases is given below: Case 1: \ud835\udf03 = 0\u00b0 The WFW is computed for a constant orientation of \ud835\udf03 = 0\u00b0 for different grid sizes as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001363_s12206-020-0734-y-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001363_s12206-020-0734-y-Figure1-1.png",
        "caption": "Fig. 1. (a) Major IPMSM motor components; (b) detailed structural view of rotor components.",
        "texts": [
            " The air circulation is intended to facilitate the air-flow inside the motor and consequently to lower the magnet temperature of the rotor. In general, the shaft rotation of an electric motor only weakly affects both radial and vertical flows in the motor, owing to an extremely thin boundary layer close to the rotating shaft in the radial direction and a negligible fluid flow in the axial direction. Therefore, protrusion-shaped flow inducers were introduced in the present work to steer local advection through large hollows (i.e., ventilation holes), as illustrated in Fig. 1, and the local convective heat transfer rate along the vertical circulation flow was evaluated by in-depth thermofluidic analysis. Subsequently, the radial heat transfer rate across the airgap between the rotor and stator was also analyzed, using the TaylorCouette flow paradigm. A 40 kW interior permanent magnet synchronous motor (IPMSM) for electric vehicular applications was adopted as an analytical model in this study; the motor is depicted in Fig. 2. Permanent magnets are installed in the rotor of the IPM motor, which produces magnetic fields for turning motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure2-1.png",
        "caption": "Fig. 2. Prototype STPSPM machines. (a) 10-pole/12-slot. (b) 10-pole/9-slot.",
        "texts": [
            " For DTPPM machines with each winding group being rotationally asymmetric, postdemagnetization UMF should be checked even without eccentricity. Finally, the experiments are conducted to validate theoretical analysis. Since surface-mounted PM (SPM) machines usually suffer from higher demagnetization risks than interior PM (IPM) machines [15], SPM machines are investigated as an example. A 10-pole/12-slot and a 10-pole/9-slot single three-phase SPM (STPSPM) machines are selected for investigation. Finite element (FE) method is used and FE models are shown in Fig. 2. The parameters are listed in TABLE I. The 10-pole/12-slot prototype machine can also be investigated as dual three-phase SPM (DTPSPM) machine, only demanding changes in the winding configuration, which will be detailed later. The demagnetization curve of N45H is shown in Fig. 3. The demagnetization mechanism is explained for a certain point on the PM as follows. Under open-circuit condition, the PM working point lies on the point N. When working current is applied, the armature field can move the PM working point along the linear part of demagnetization curve from the point N to the point M",
            " Hence, there is no UMF when two winding groups are normally supplied without considering demagnetization and eccentricity faults. However, since the two winding groups are usually supplied by two isolated converters, it is almost impossible for 3PSC to happen in two winding groups simultaneously. Therefore, 3PSC is usually assumed to happen in only one winding group whilst the remaining one is assumed to be normally supplied [19]-[21]. Under such condition, the post-demagnetization UMF heavily depends on the dual three-phase winding configuration. The 10-pole/12slot SPM machines shown in Fig. 2(a) yet with three typical dual three-phase winding configurations are investigated as an example, which are shown in Fig. 8. The 3PSC currents in the three DTPSPM machines are shown in Fig. 9 and the demagnetization distributions are 0 20 40 60 80 0 1 2 3 4 5 6 7 8 9 10 11 12 F x ( N ) Harmonic order Fx-before Fx-after -100 -50 0 50 100 0 60 120 180 240 300 360 F y (N ) Rotor position (Mech. Deg.) Fy-before Fy-after Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore",
            " Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. A. Prototype SPM machines In this section, experiments are conducted to validate the numerical predictions. STPSPM machines are manufactured and tested, which are shown in Fig. 25. The parameters are the same with the 10-pole/12-slot STPSPM machines shown in Fig. 2(a). However, one of the rotors lacks two poles, which can be regarded as being symmetrically demagnetized, whilst the other one lacks one pole, which can be regarded as being unsymmetrically demagnetized. Firstly, the back-EMF waveforms for two STPSPM machines are measured. As shown in Fig. 26, the measured back-EMF generally matches well with the FE predictions. In addition, the cogging torque waveforms for two STPSPM machines are also measured, where the method in [26] is employed. As shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure6-1.png",
        "caption": "Fig. 6. Rigid domain allocated to an artificial domain which will not be part of the printed outcome. The Fixed constraint is applied to a domain which is excluded from the optimization. All other boundaries are free.",
        "texts": [
            " The overall, time-independent mechanical problem consists of solving \u2207 \u00b7 S = Fv, (1) where Fv is an external volume force - in our case zero - and S is the stress tensor S = D0T. (2) In the above, D0 denotes the material stiffness tensor and T is the linear strain tensor with components given by Tij = ( \u2202ui \u2202xj + \u2202uj \u2202xi ) . (3) In cartesian coordinates, i, j \u2208 {x, y, z} and u is the material deflection. Equation (1) is subjected to boundary conditions, being either S = 0 (free boundary) or u = 0 (fixed boundary). These respective conditions are indicated on the geometry shown in Figure 6. Since the application requires some boundaries to be maintained in place, i.e. the motor enclosure and the back end of the non-conductive holder, meant to press on the battery pack, the optimization is performed on the domain shown in Figure 7. In addition to the deformable material, we have modelled the battery pressure by a rigid domain subjected to a prescribed displacement of 0.5 mm while being free to rotate. Solving eq. (3) above gives the classical solutions to mechanical problems without any optimization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002839_1045389x211014571-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002839_1045389x211014571-Figure1-1.png",
        "caption": "Figure 1. Deformation of a material line element dX to a spatial line element dx.",
        "texts": [
            " This causes a deformation to the material. To enhance the interpretation of the presented results, in this section the kinematics and the necessary variables, based on a continuum mechanical perspective, are established. Due to the applied load in the experiment the material undergoes a deformation. The deformation can herein be described as the relation of certain material particle coordinates X in the undeformed configuration to the coordinates x of the deformed configuration (Eringen, 1980), see Figure 1. The particle performs a movement x(X, t) from the initial position X at time t= 0 to the current position x at t.0. The displacement u can be described as the difference of the two coordinates u= x(X, t) X: \u00f01\u00de A widly used measure of deformation is the deformation gradient F (Holzapfel, 2002), transforming an undeformed line element dX into a deformed line element dx via F(X, t)= \u2202x \u2202X : \u00f02\u00de By means of a polar decomposition, the deformation gradient can be subdivided into its stretching part U and its rotatory part R F=R U: \u00f03\u00de Due to the fact, that a rigid body rotation of a material particle does not lead to a deformation, the rigid body rotation has to be separated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000041_012009-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000041_012009-Figure9-1.png",
        "caption": "Figure 9: Electrostatic field impact on UAV actuation section at (a) azimuth angle \u03c9 = 120 degrees (b) azimuth angle \u03c9 = 240 degrees",
        "texts": [],
        "surrounding_texts": [
            "To avoid the corona discharge interference, a shielding solution is recommended. However, navigating a floating conductor inside the valve hall can introduce the risk of changing air breakdown characteristics inside the valve hall. Hence, an impulse test was conducted to evaluate the maximum change of the air breakdown voltage in the presence of a floating conductor. The impulse DC test setup is shown in Fig. 10, where the drone model is tethered in the midway between two rod electrodes. The drone model is a cross aluminum bar which has the 16th Deep Sea Offshore Wind R&D conference IOP Conf. Series: Journal of Physics: Conf. Series 1356 (2019) 012009 IOP Publishing doi:10.1088/1742-6596/1356/1/012009 same dimensions of a typical drone (0.5 m X 0.5 m) and thickness of 0.01 m, which represents the worst-case shielding scenario. The upper rod electrode is connected to a HVDC power supply, while the lower electrode is connected to ground. The two electrodes are separated by 1.7 m air gap distance. This distance is selected since the vertical distance between HVDC valve hall racks is from 1 m to 1.7 m [19]. In negative impulse test, the breakdown voltage between the two electrodes is found to be -1.158 MV and -1.126 MV in both absence and presence of the drone model, respectively. In positive impulse test, the breakdown voltage between the two electrodes is found to be 1.021 MV and 1.002 MV in absence and presence of the drone model, respectively. The experimental results agree with the results introduced in [18], in case of negative air breakdown. In positive air breakdown, the aluminum cross bar floating conductor decreases slightly the breakdown voltage compared to a corresponding increase in case of using spherical conductor. Within the realms of these tests, there has been no change to the breakdown voltage but further investigations would need to be done to confirm other parameters and probabilities. This is considered as a preliminary result, which needs further testing using real shielded drone 16th Deep Sea Offshore Wind R&D conference IOP Conf. Series: Journal of Physics: Conf. Series 1356 (2019) 012009 IOP Publishing doi:10.1088/1742-6596/1356/1/012009 at different orientations and distances from energized electrodes."
        ]
    },
    {
        "image_filename": "designv11_63_0003282_j.ijnonlinmec.2021.103808-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003282_j.ijnonlinmec.2021.103808-Figure3-1.png",
        "caption": "Fig. 3. Free body diagram of leading shoe during contact with the brake drum.",
        "texts": [
            "\ud835\udc51 ) +\ud835\udc3e\ud835\udc60(\ud835\udf03\ud835\udc53 (\ud835\udc61) \u2212 \ud835\udf03\ud835\udc51 (\ud835\udc61)) = 0, (17) \ud835\udc51 ?\u0308?\ud835\udc51 + \ud835\udc36\ud835\udc60(?\u0307?\ud835\udc51 \u2212 ?\u0307?\ud835\udc53 ) +\ud835\udc3e\ud835\udc60(\ud835\udf03\ud835\udc51 (\ud835\udc61) \u2212 \ud835\udf03\ud835\udc53 (\ud835\udc61)) = \ud835\udf0f\ud835\udc51 (\ud835\udc61). (18) he frictional torque on the drum, \ud835\udf0f\ud835\udc51 (\ud835\udc61) is given by the following where \ud835\udc53 \ud835\udc59 (\ud835\udc61) and \ud835\udc39 \ud835\udc53 \ud835\udc5f (\ud835\udc61) are the frictional forces between the drum and the hoes. \ud835\udc51 (\ud835\udc61) = \u2212(\ud835\udc39 \ud835\udc53 \ud835\udc59 (\ud835\udc61) + \ud835\udc39 \ud835\udc53 \ud835\udc5f (\ud835\udc61))\ud835\udc45\ud835\udc51 . (19) he frictional forces were formulated, as shown in Eq. (20), where \ud835\udf07(\ud835\udc61) s the time-varying coefficient of friction, while \ud835\udc39 \ud835\udc5b(\ud835\udc61) is the normal force etween the shoe and the drum. \ud835\udc53 \ud835\udc59 (\ud835\udc61) = \ud835\udf07\ud835\udc59(\ud835\udc61)\ud835\udc39 \ud835\udc5b \ud835\udc59 (\ud835\udc61), (20a) \ud835\udc53 \ud835\udc5f (\ud835\udc61) = \ud835\udf07\ud835\udc5f(\ud835\udc61)\ud835\udc39 \ud835\udc5b \ud835\udc5f (\ud835\udc61). (20b) Fig. 3 shows the free body diagram of the leading brake shoe in a eneral contact position with the drum making an angle \ud835\udf03\ud835\udc59 from the ero position. The normal and the frictional forces are acting at the oint of contact (\ud835\udc43\ud835\udc59). Dynamic moment balance of the brake shoe about he pivot point \ud835\udc42\ud835\udc59 leads to the following equation of motion, where \ud835\udc3c\ud835\udc59 s the mass moment of inertia of the leading shoe about \ud835\udc42\ud835\udc59. \ud835\udc59 ?\u0308?\ud835\udc59 + \ud835\udc39 \ud835\udc5b \ud835\udc59 \ud835\udc40 \ud835\udc5b \ud835\udc59 \u2212 \ud835\udc39 \ud835\udc53 \ud835\udc59 \ud835\udc40 \ud835\udc53 \ud835\udc59 \u2212 \ud835\udc39 \ud835\udc60 \ud835\udc59 \ud835\udc40 \ud835\udc60 \ud835\udc59 \u2212 \ud835\udc39\ud835\udc64 \ud835\udc59 \ud835\udc40\ud835\udc64 \ud835\udc59 = \ud835\udc39 \ud835\udc52 \ud835\udc59 \ud835\udc40 \ud835\udc52 \ud835\udc59 . (21) ere, \ud835\udc40 represents the time-dependent moment arm of the correspondng forces about the pivot point, as shown in Fig. 3. Similarly, for the railing shoe, the equation of motion was derived as shown, \ud835\udc5f?\u0308?\ud835\udc5f + \ud835\udc39 \ud835\udc5b \ud835\udc5f \ud835\udc40 \ud835\udc5b \ud835\udc5f + \ud835\udc39 \ud835\udc53 \ud835\udc5f \ud835\udc40 \ud835\udc53 \ud835\udc5f \u2212 \ud835\udc39 \ud835\udc60 \ud835\udc5f \ud835\udc40 \ud835\udc60 \ud835\udc5f \u2212 \ud835\udc39\ud835\udc64 \ud835\udc5f \ud835\udc40\ud835\udc64 \ud835\udc5f = \ud835\udc39 \ud835\udc52 \ud835\udc5f \ud835\udc40 \ud835\udc52 \ud835\udc5f . (22) here \ud835\udc3c\ud835\udc5f is the mass moment of inertia of the trailing shoe about \ud835\udc42\ud835\udc5f. ased on the contact theory [30], \ud835\udc39 \ud835\udc5b is given by: \ud835\udc5b(\ud835\udc61) = \ud835\udc58\ud835\udc50\ud835\udf16\ud835\udc56(\ud835\udc61) + \ud835\udc39 \ud835\udc51 (\ud835\udc61). (23) ere, the linear contact stiffness (\ud835\udc58\ud835\udc50) was defined for a line contact ased on Hertzian contact theory [30] as shown below \ud835\udc50 = \ud835\udf0b 4 \ud835\udc38\ud835\udc52\ud835\udc5e\ud835\udc59 \ud835\udc50 . (24) where \ud835\udc38\ud835\udc52\ud835\udc5e and \ud835\udc59\ud835\udc50 are the equivalent Young\u2019s modulus and length of contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002137_j.mechatronics.2020.102482-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002137_j.mechatronics.2020.102482-Figure5-1.png",
        "caption": "Fig. 5. Schematic of AMBS System.",
        "texts": [
            " Furthermore, the experimental set-up and GTDF design process are discussed followed by the illustration of the experimental results. 5.1. System description and experimental set-up AMBS consists of a series of electromagnets arranged in a ring to create a contact-free rotor bearing. By eliminating mechanical contact between the rotor and housing, AMBS enables friction-less operation of the rotor. AMBS creates friction-less support of the rotor and eliminates contact bearing\u2019s thermal expansions and mechanical wears. The experimental system shown in Fig. 5 consists of eight electromagnet amplifiers to inject currents to the coils. The nonlinear electromagnetic forces caused by the coils\u2019 currents levitate the rotor. To make AMBS more linear, bias currents are added to opposing coils. Four sensors are located at two ends of the rotor to measure its position with respect to the housing. Also, a LABVIEW Real-Time target installed on the computer, sampling at 10 kHz is used for real-time digital control. Since the axes of the sensors are not co-located with those of the actuators, the system has strong coupling effects"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002629_012013-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002629_012013-Figure3-1.png",
        "caption": "Figure 3. End cover flange to motor housing",
        "texts": [
            " To replace rotor coil, a point mass load is added to the assembly. Rotor shaft Z - translation has been constrained. Also no separation added (Face-Face) between contact faces. In this pair, one part is acting as contact body and other one is target body.Motor housing base is constrained as a fixed support, i.e., there is no translation movement in X, Y and Z directions. To replace rotor coil, a point mass load is added to the assembly. Rotor shaft Z - translation has been constrained by applying rotational velocity from 200 to 1200 rpm. Figure 3, 4,5 and 6 shows the bonded contact for the electrical motor. Industry 4.0 Technologies in Civil and Mechanical Engineering (ICI4TCME 2020) IOP Conf. Series: Materials Science and Engineering 1112 (2021) 012013 IOP Publishing doi:10.1088/1757-899X/1112/1/012013 Modal Analysis is mostly dispensed to seek out natural frequency of the system and completely different mode shapes at different Eigen values [12,14]. In this work, totally different modes and Eigen values square measure calculates and premeditated between mode shapes versus natural frequencies [15,20, Industry 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002703_j.optlastec.2021.107100-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002703_j.optlastec.2021.107100-Figure4-1.png",
        "caption": "Fig. 4. Precutting samples of 304SS with the thicknesses of 1 mm and 2 mm.",
        "texts": [
            " Then, the milled plates are ultrasonically cleaned in the acetone and then dried for processing. X. Lin et al. Optics and Laser Technology 141 (2021) 107100 Firstly, as previous discussion, 1 mm and 2 mm 304SS metal plates are successfully precut by DLS. A set of standard parameters for diode laser cutting is used. The laser power is 1 kW, and the cutting speed is 5 m/min. During the cutting processing, N2 with a flow rate of 20 L/min is used as a shielding gas. These parameters are optimized for minimizing burrs. Cutting samples are shown in Fig. 4. Then, we select eligible precut samples for the welding. Many welding experiments are carried out according to the above processing procedure. In order to determine the influence of laser welding parameters on the quality of TWBs, these welding experiments are performed at a variation of laser power, P, from 1.5 to 2 kW and the welding Table 1 Wavelength channels of three submodules used in DLS. Channel 1 Channel 2 Channel 3 Channel 4 Channel 5 Module 1 (91x nm) 913.1 914.5 916.0 917.6 919.0 Module 2 (94x nm) 941"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003282_j.ijnonlinmec.2021.103808-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003282_j.ijnonlinmec.2021.103808-Figure2-1.png",
        "caption": "Fig. 2. Side view of the drum brake showing the different orientations of the leading shoe. Note: In the figure, only leading shoe is shown. (a) Zero state. (b) Kinematic contact state. (c) General contact state (with a magnified indentation for clarity).",
        "texts": [
            " Thus, the possibility of tribological phenomena (such as adhesion, fretting, and fatigue between brake drum and brake shoe) are very low and hence not considered. Further, the surface evolution has not been considered part of the study, as the surface parameters do not change considerably during a braking event. 3. Development of the analytical model 3.1. Derivation of the kinematic relationship It is essential to establish the kinematic relationship before deriving the dynamic equations of motion, as the drum brake geometry is complex. Various states of the system are depicted in Fig. 2; specifically, the zero state (as defined earlier) is shown in Fig. 2(a). First, the contact point between the brake shoe and the brake drum was obtained kinematically as the intersection point of the circular brake drum, and the brake shoe (circular arc) rotated about the pivot point. 3.1.1. Determination of the location of contact point Fig. 2(b) shows the drum brake system in the kinematic contact state (represented by the superscript \ud835\udc58), where the brake drum and shoe kinematically touch each other. Geometrically, \ud835\udc3a\ud835\udc59 and \ud835\udc3a\ud835\udc5f does not lie along the \ud835\udc65-axis (as shown in Fig. 2(b)) during the kinematic state as opposed to the zero state. The contact point on the brake drum was defined as \ud835\udc43 \ud835\udc58 \ud835\udc51 , while the same on the leading shoe was defined as \ud835\udc43 \ud835\udc58 \ud835\udc59 . As seen, \ud835\udc43 \ud835\udc58 \ud835\udc51 and \ud835\udc43 \ud835\udc58 \ud835\udc59 coincide at the kinematic contact state, and the ngle made by the leading shoe is \ud835\udf03\ud835\udc58\ud835\udc59 . The procedure for calculating \ud835\udf03\ud835\udc58\ud835\udc59 is iven below, where the global coordinates of \ud835\udc43 \ud835\udc58 \ud835\udc59 are given by Eqs. (2a) nd (2b). \ud835\udc58 \ud835\udc59 | | |\ud835\udc65 = \ud835\udc45\ud835\udc59 cos(\ud835\udefc0\ud835\udc59 ) cos(\ud835\udf03 \ud835\udc58 \ud835\udc59 ) \u2212 \ud835\udc45\ud835\udc59 sin(\ud835\udefc0\ud835\udc59 ) sin(\ud835\udf03 \ud835\udc58 \ud835\udc59 ) + | \u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d7\ud835\udc3a\ud835\udc51\ud835\udc3a\ud835\udc59|\ud835\udc65, (2a) \ud835\udc58 \ud835\udc59 | | |\ud835\udc66 = \ud835\udc45\ud835\udc59 cos(\ud835\udefc0\ud835\udc59 ) sin(\ud835\udf03 \ud835\udc58 \ud835\udc59 ) + \ud835\udc45\ud835\udc59 sin(\ud835\udefc0\ud835\udc59 ) cos(\ud835\udf03 \ud835\udc58 \ud835\udc59 ) + | \u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d6\u20d7\ud835\udc3a\ud835\udc51\ud835\udc3a\ud835\udc59|\ud835\udc66",
            " (7) yields two roots corresponding to the two contact points (when two circles intersect). However, the kinematic contact point is the one when both the obtained roots are real and repeated. Thus, by applying the condition for repeated roots (\ud835\udc352 \u2212 a A w a i a a g \ud835\udc43 \ud835\udc43 A a \ud835\udf16 F \ud835\udf16 \ud835\udf16 H w \ud835\udc5b \ud835\udf16 \ud835\udf16 \ud835\udc3c T \ud835\udc39 s \ud835\udf0f T i b \ud835\udc39 \ud835\udc39 g z p t i \ud835\udc3c H i t \ud835\udc3c w B \ud835\udc39 H b \ud835\udc58 4\ud835\udc34\ud835\udc37 = 0), for a given drum brake geometry, \ud835\udf03\ud835\udc58\ud835\udc59 was obtained, and hence \ud835\udefc0\ud835\udc59 and \ud835\udefc\ud835\udc58\ud835\udc59 (using Eqs. (3) and (5)). 3.1.2. Calculation of the indentation and its time derivative Fig. 2(c) shows the drum brake system with the leading shoe in a general rotated state with finite indentation (\ud835\udf16\ud835\udc59(\ud835\udc61)) when \ud835\udc39 \ud835\udc52 \ud835\udc59 (\ud835\udc61) is applied. Distance between the contact points on the brake drum (\ud835\udc43 \ud835\udc58 \ud835\udc51 ) nd the brake shoe (\ud835\udc43\ud835\udc59) is the indentation \ud835\udf16(\ud835\udc61) as shown in Fig. 2(c). moving coordinate system (\ud835\udc56, \ud835\udc57) was defined, fixed to the brake shoe ith its origin at \ud835\udc43\ud835\udc59 where \ud835\udc56 is along the normal direction (directed way from \ud835\udc3a\ud835\udc51) and while \ud835\udc57 is along the tangential direction as shown n Fig. 2(a). The instantaneous value of indentation and its components long normal and tangential directions (\ud835\udf16\ud835\udc56\ud835\udc59(\ud835\udc61) and \ud835\udf16\ud835\udc57\ud835\udc59 (\ud835\udc61)) were calculated s given below. General expressions for the coordinates of \ud835\udc43\ud835\udc59(\ud835\udc61), for a iven \ud835\udf03\ud835\udc59(t) are, \ud835\udc59(\ud835\udc61) | | |\ud835\udc65 = \ud835\udc45\ud835\udc59 cos(\ud835\udefc0\ud835\udc59 ) cos(\ud835\udf03\ud835\udc59(\ud835\udc61)) \u2212 \ud835\udc45\ud835\udc59 sin(\ud835\udefc0\ud835\udc59 ) sin(\ud835\udf03\ud835\udc59(\ud835\udc61)) + (\ud835\udc5d\ud835\udc59 \u2212 \ud835\udc52\ud835\udc59) cos(\ud835\udf03\ud835\udc59(\ud835\udc61)) \u2212 \ud835\udc5e\ud835\udc59 sin(\ud835\udf03\ud835\udc59(\ud835\udc61)) \u2212 \ud835\udc5d\ud835\udc59 , (9a) \ud835\udc59(\ud835\udc61) | | |\ud835\udc66 = \ud835\udc45\ud835\udc59 cos(\ud835\udefc0\ud835\udc59 ) sin(\ud835\udf03\ud835\udc59(\ud835\udc61)) + \ud835\udc45\ud835\udc59 sin(\ud835\udefc0\ud835\udc59 ) cos(\ud835\udf03\ud835\udc59(\ud835\udc61)) + (\ud835\udc5d\ud835\udc59 \u2212 \ud835\udc52\ud835\udc59) sin(\ud835\udf03\ud835\udc59(\ud835\udc61)) + \ud835\udc5e\ud835\udc59 cos(\ud835\udf03\ud835\udc59(\ud835\udc61)) \u2212 \ud835\udc5e\ud835\udc59 . (9b) s per the definition of (\ud835\udf16\ud835\udc59(\ud835\udc61)) and considering its sign, it was calculated s, \ud835\udc59(\ud835\udc61) = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 + \u221a (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc65 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc65 )2 + (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc66 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc66 )2 if \u221a (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc65 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc65 )2 + (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc66 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc66 )2 > \ud835\udc45\ud835\udc51 \u2212 \u221a (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc65 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc65 )2 + (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc66 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc66 )2 if \u221a (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc65 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc65 )2 + (\ud835\udc43\ud835\udc59(\ud835\udc61) | | |\ud835\udc66 \u2212 \ud835\udc43 \ud835\udc58 \ud835\udc51 | | |\ud835\udc66 )2 \u2264 \ud835\udc45\ud835\udc51 (10) urthermore, \ud835\udf16\ud835\udc59(\ud835\udc61) was resolved along \ud835\udc56 and \ud835\udc57 as, \ud835\udc56 \ud835\udc59(\ud835\udc61) = \ud835\udf16\ud835\udc59(\ud835\udc61) cos(\ud835\udef7\ud835\udc59(\ud835\udc61)), (11a) \ud835\udc57 \ud835\udc59 (\ud835\udc61) = \ud835\udf16\ud835\udc59(\ud835\udc61) sin(\ud835\udef7\ud835\udc59(\ud835\udc61))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002976_j.mechatronics.2021.102584-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002976_j.mechatronics.2021.102584-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of the tendon-sheath mechanism.",
        "texts": [
            " Finally, conclusions and future work are discussed in Section 6. The tendon-sheath mechanism has severe hysteresis transmission characteristics due to the varied friction between the tendon and the sheath. In addition, the elastic springs further complicate the nonlinearity of the system and bring challenges to the control of the tendonsheath artificial muscle. In this section, the tendon-sheath artificial muscle modeling is carried out based on a tendon-sheath transmission model. The tendon-sheath transmission mechanism, as Fig. 2 shows, has an outer sheath and an inner tendon. The outer sheath has a slender body and a smooth inner wall, and is flexible and axially incompressible. The inner tendon is usually a wire rope with a diameter slightly smaller than the inner diameter of the sheath. The tendon can transfer force up to thousands of Newton when its diameter increases (usually less than 3 mm). In practical applications, the two ends of the outer sheath are fixed, and the sheath body is bent as needed. The inner tendon moves in the sheath to transmit force and displacement from one end to another"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002578_j.jclepro.2021.126900-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002578_j.jclepro.2021.126900-Figure1-1.png",
        "caption": "Fig. 1. A) Structural schematisation of the whole stator (OX \u00bc 300 mm and OY \u00bc 600 mm) with details of the single sector assembly (red dashed square); B) analysed three-bladed sector.",
        "texts": [
            " The FAB process is characterised by the absence of processing waste due to the high degree of automation and repeatability. In summary, the developed FAB technology proves to be a much cheaper and less environmentally impactful process than the traditional technology. The present paper describes scientific research carried out on aero-engine turbine blades. In particular, the experimental analysis focused on evaluating the final surface quality of a sector of the stator constituted by three blades, as shown in Fig. 1. This component is made of a nickel-based superalloy through the lost wax casting process, also known as \u201cinvestment casting\u201d (Torres-Carrillo et al., 2020). The choice of the constructionmaterial is due to the outstanding performance of these alloys in hightemperature environments; this property makes them the most suitablematerials for high-severity applications (Vilaro et al., 2012). The whole stator consists of 20 of these bladed sectors, each with an angle of 18 . The external radius of the whole stator is approximately 600 mm, and the internal radius is approximately 300 mm. The weight of the complete stator is approximately 34.26 kg before the final surface finishing processing. As shown in Fig. 1, the geometry of the sector is very complex. This makes surface finishing operations with conventional systems long and difficult. For this reason, the process is carried out by hand by specialised operators who must follow a long training period. S. Venettacci, G.S. Ponticelli and S. Guarino Journal of Cleaner Production 299 (2021) 126900 Investment casting, or lost wax casting, is a metal-forming technique that consists of eight main steps (Kalpakjian and Schmid, 2014), summarised in Fig. 2: i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001914_asyu50717.2020.9259819-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001914_asyu50717.2020.9259819-Figure7-1.png",
        "caption": "Fig. 7. The proposed gear early fault diagnosis test setup.",
        "texts": [
            " The single tooth stiffness of the gears is calculated for both healthy and cracked gears, and then the time-varying mesh stiffness is calculated by using the single tooth stiffness results based on the previous study [12] as given in Fig.6. Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 24,2020 at 15:25:26 UTC from IEEE Xplore. Restrictions apply. This paper proposed a test setup to detect early gear faults in railway gearboxes as future work for contributing to railbased transportation system safety, reliability, and security in a medium to long term perspective. Fig.7 shows the threedimensional CAD design of the gear early fault diagnosis test setup. The input of a single-stage gearbox is connected to the electric motor and its output to the magnetic brake. A motor drive is used to fix the motor at different operating speeds. Motor and brake connections are provided with rigid couplings to facilitate the removal and installation of different gears during the experiments. Besides, a torque meter is used to measure torque between the gearbox and motor. A singlestage gearbox design is preferred to assemble and disassemble different gears easily"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000720_j.engstruct.2020.110416-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000720_j.engstruct.2020.110416-Figure9-1.png",
        "caption": "Fig. 9. The bladed disk model in finite element analysis.",
        "texts": [
            "0 kg and a standard deviation \u03c3 = 0.01 kg. However, the whole simulation time based on sensitivity analysis is about 25 s CPU time while the exact simulation of 5 million eigensolutions takes about 32 h CPU time under Matlab platform on an Intel\u00ae Core\u2122 i7-4770 CPU@3.40 GHz desktop computer. To further validate the proposed eigensensitivity analysis, finite element modeling of a bladed disk with 24 blades has been carried out. The geometrical and material properties of the bladed disk are shown in Fig. 9. Commercial finite element software ANSYS 19.2 was used and typical vibration mode shapes for the tuned case are shown in Fig. 10. The same structure was carefully fabricated and tested using Laser Dopler Vibraometer (LDV) based modal testing and analysis system and some typical measured vibration modes are shown in Fig. 11. The test results have shown that the fabricated bladed disk is very close to a tuned condition since the doubly repeated modes all appear at a single frequencies as observed during the vibration test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001309_acc45564.2020.9147355-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001309_acc45564.2020.9147355-Figure4-1.png",
        "caption": "Figure 4. Definition of coordinate frames and free-body-diagram of the payload, where point P is the CoG, B and I denote the body-fixed and inertial frame, respectively.",
        "texts": [
            " Based on (1) and (2), the dynamics of the system in the UKF can be expressed as x\u0307F = f1 (xF , uF ) + w1 yF = h1(xF ) + v1, where f1\u2208 R12 is a nonlinear function that characterizes the dynamics of the follower, uF = [ 0, 0, TF ]T \u2208 R3 represents the control input, w1\u2208 R12 and v1\u2208 R2 are defined as the process and measurement noises, respectively, yF\u2208 R6 is the measurement, xF is the follower state defined as xF = [ pT F vT F aT F FT F ]T , (3) and h1(xF )\u2208 R6 is a measurement model defined as h1(xF ) = [ pT F aT F ]T , (4) where pF is the follower\u2019s position measured by motion capture system2, and aF is the follower\u2019s acceleration measured by the IMU from the flight controller. In this section, the external force of the leader exerted to the payload is estimated, which is FL as shown in Fig. 4. Since the external force is applied by the cable connecting to the payload, the dynamics of the payload is modeled and analyzed. To facilitate the subsequent analysis, a assumption is made as follows: Assumption 3. The payload is a rigid body. An IMU is attached to the payload at point C in Fig. 4 for measuring the acceleration and angular velocity. The dynamics of the payload is required for state estimation in the second UKF and is described as aBp = aBc \u2212 \u03b1B p \u00d7 rcp + \u03c9B p \u00d7 (\u03c9B p \u00d7 rcp) (5) FL = mpRpa B p \u2212 FF +mpg, (6) where (\u00b7)B denotes the vector expressed in the body-fixed frame on the payload, FF \u2208 R3 is the estimate outputted from the first UKF developed in Section III-A, ap \u2208 R3 is the acceleration of point P, ac \u2208 R3 denotes the acceleration of point C, which is measured from the IMU attached to the payload, \u03b1p \u2208 R3 denotes the angular acceleration, rcp \u2208 R3 is the vector from point C to the center of mass P, \u03c9p \u2208 R3 denotes the angular velocity, Rp \u2208 SO(3) is the rotation 2In outdoor environment, the motion capture system can be replaced by GPS"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure52.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure52.3-1.png",
        "caption": "Fig. 52.3 (1) Engine block; (2) Support Frame; (3) Blade; (4) Flange; (5) Rotor shaft casing; (6); Pulley; (7) Rotor shaft; (8) Bearing; (9) Pulley belt; (10) Pulley; (11) Vertical lifting arm; (12) Bracket; (13) Shaft connecting to the power tiller; (14) Arm; (15) Arm locking plate with pin",
        "texts": [
            " The other end of the arm is attached to the shaft extending from the power tiller by means of a bracket. An existing power tiller is considered as another sub-unit required to provide the forward and reverse speed of the whole system. The power required for cutting the soil to the rotary cutting unit will be provided by the engine mounted on top of the rotary cutting unit. Each individual partwere created first in \u201cPartDesign\u201dworkbench and later assembled on \u201cAssembly Design\u201d workbench. The assembled design of the system was simulated on \u201cDMU Kinematics\u201d workbench (Fig. 52.3). It consists of six blades made of tool steel of 5 mm thickness mounted on three flanges mounted on the shaft casing pipe. The flanges are in the shape of a circular disc having six cut out slots along the circumference in order to house the blades (Fig. 52.4a). The support frame unit consists of the main frame upon which the engine will be mounted. The bearings and the shafts to be connected to the rotary cutting unit are mounted to this frame as well. A plate in the shape of an arc which provides locking positions of the arm, and the support frame is also fixed to this frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000250_tcsii.2019.2957543-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000250_tcsii.2019.2957543-Figure1-1.png",
        "caption": "Fig. 1. Dimensions of the designed ESPAR antenna.",
        "texts": [
            " In this letter, we demonstrate how single-anchor indoor localization concept can be advanced in terms of achievable accuracy by using a base station equipped with the designed ESPAR antenna and by employing both, directional main beam and narrow minimum, radiation patterns. Measurements of the prototype performed in a real-world scenario indicate that even for the proposed simple localization algorithm based on signal strength recorded in the BS, one can easily obtain levels of accuracy similar to the one for the most sophisticated algorithm presented in [1]. The proposed design comprises 12 passive elements ESPAR antenna (Fig. 1) with one active monopole in the center of the 1536-1225 \u00a9 2015 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. ground plane being a top layer of the printed circuit board (PCB) base. The active element is fed by an SMA connector, while the parasitic elements can be connected to the ground or opened by the single-pole, single-throw (SPST) switches connected to the end of each of them at the bottom layer of the structure. Parasitic elements connected to the ground are referred to as reflectors because they reflect energy, while the opened elements are directors as the electromagnetic wave can pass through them. All switches are controlled by an external microcontroller, hence the actual configuration of the antenna can be denoted by the steering vector , where denotes the state of each parasitic element in Fig. 1: for th parasitic element connected to the ground and 1 for opened. The antenna was designed and simulated in FEKO electromagnetic simulation software tool. The antenna design is based on those proposed in [4] and [5] and employs 1.55-mm-height FR4 laminate with top-layer metallization. In [4], the number of 12 parasitic elements was proven to be optimal with respect to the narrowest main lobe and the lowest backward radiation at the center frequency equal to 2.4 GHz, hence the same starting configuration was chosen, and then the antenna was optimized to obtain various radiation patterns (see Section II-B) and satisfactory input impedance matching for all parasitic elements configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001865_012013-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001865_012013-Figure5-1.png",
        "caption": "Figure 5. Sketch of the device for running-in",
        "texts": [
            " As a tool, it is proposed to use a roller with a toroidal deforming surface [18], which will provide the possibility of rolling the cycloidal helical surface of rotors with different geometric parameters on standard CNC equipment. To implement the method of rolling a complex profile surface on a multi-purpose CNC machine, it is necessary to develop a tool that allows running on metal-cutting equipment [19] [20]. As a tool for processing, it is proposed to use a device containing a roller with a toroidal working surface, a housing that acts as a guide and accommodates an elastic element, and a tool cone for installation in the machine spindle by an automatic tool change system (figure 5). Creating the necessary pressure in the contact zone of the roller and the workpiece is carried out by deforming the elastic element inside the device body by a certain amount. The constant value of the deformation Modern power engineering (MPMB 2020) IOP Conf. Series: Materials Science and Engineering 963 (2020) 012013 IOP Publishing doi:10.1088/1757-899X/963/1/012013 is provided by controlling the processing process by means of the CNC. Moreover, given that the stiffness of the technological system can differ significantly at different points, the value of the working deformation of the elastic element should be an order of magnitude greater than the value of the deformations of the technological system under the action of the rolling force [21]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure13.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure13.5-1.png",
        "caption": "Fig. 13.5 Heuristics extraction and analysis process",
        "texts": [
            "5). A total of 650 furniture were selected for analysis. The details of the existing furniture were analyzed based on the form and structure and special features. The formand structurewere further deconstructed into bar, body, and surface [41] and the orientation and dominance of the elementwere also analyzed [42]. After the analysis, existing furniture was grouped (Fig. 13.4a) according to the similarity, as suggested in the K-J method (Fig. 13.4b) [43, 44]. Analysis and heuristics extraction (Fig. 13.5) of the existing furniture has been done, as suggested by Yilmaz et al. [22]. The success of the heuristics extraction process and analysis was evaluated with the inter-rater agreement between three coders, which measures the degree to which they agree to each other. In this process, one Professor, one Assistant Professor and oneMaster of Design Student participated. The initial inter-rater reliability was 90%. The disagreement between the coders was resolved through discussions and 100% reliability was achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure6-1.png",
        "caption": "Figure 6. First and second eigenmode (51.16Hz and 78.29Hz). The scale for the magnitude of the eigenmodes is the same for all figures which follow.",
        "texts": [],
        "surrounding_texts": [
            "Composite materials, finite element analysis, automotive engineering, vibrational behavior, design optimization Date received: 21 March 2021; accepted: 5 May 2021"
        ]
    },
    {
        "image_filename": "designv11_63_0000880_j.mechmachtheory.2020.103900-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000880_j.mechmachtheory.2020.103900-Figure11-1.png",
        "caption": "Fig. 11. Visualization of the force distributions \u02c6 e1 and \u02c6 e2 . The color of the actuators indicates the value of active force: white indicates 1 N, black \u22121 N , and gray 0 N. The arrows around the joints indicate the directions of the control torques produced by the individual muscle-like actuators.",
        "texts": [
            " m 1 m 2 m 3 m 4 m 5 m 6 Total \u03c4 1 0 0 0 0 \u2212r 1 \u2212r 1 \u22122 r 1 \u03c4 2 0 0 \u2212r 2 \u2212r 2 \u2212r 2 \u2212r 2 \u22124 r 2 K q 11 0 0 0 0 \u2212r 2 1 k 5 r 2 1 k 6 r 2 1 ( k 6 \u2212 k 5 ) K q 12 0 0 0 0 \u2212r 1 r 2 k 5 r 1 r 2 k 6 r 1 r 2 ( k 6 \u2212 k 5 ) K q 22 0 0 \u2212r 2 2 k 3 r 2 2 k 4 \u2212r 2 2 k 5 r 2 2 k 6 r 2 2 ( k 4 \u2212 k 3 + k 6 \u2212 k 5 ) 7. Discussion 7.1. Physical meaning of the new basis of U The basis given by Eq. (41) appears to depend on the arrangement of the muscle-like actuators. This section discusses the physical meaning of the new basis and how each of the basis vectors generates the control torques and the joint stiffnesses via the muscle-like actuators and their arrangement. Each basis vector \u02c6 ei in Eq. (41) represents a force distribution; the j th element of \u02c6 ei represents the active force of m j . Fig. 11 visualize the force distributions \u02c6 e1 and \u02c6 e2 , which are the basis vectors of U 1 . The color of the actuators indicates the value of active force: white indicates 1 N, black \u22121 N , and gray 0 N. The arrows around the joints indicate the directions of the control torques produced by the individual muscle-like actuators. The gray lines without arrowheads indicate that the produced control torque is zero. Fig. 11 shows that \u02c6 e1 and \u02c6 e2 activate the pair of biarticular actuators and a pair of monoarticular actuators, in which the activated flexors and extensors work in the opposite directions. We computed the contribution of each muscle-like actuator to the control torques and the joint stiffnesses when the robot arm is driven by the force distribution \u02c6 ei . This was done by the following steps: let \u02c6 ei j be a vector whose j th element is equal to that of \u02c6 ei and whose other elements are zero; the control torques can be obtained by substituting F with \u02c6 ei j in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000149_1350650119887043-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000149_1350650119887043-Figure3-1.png",
        "caption": "Figure 3. Test rig cross-section.",
        "texts": [
            " The bump foil geometry for the targeted stiffness range is arrived by following the basics of die design and existing experience in foil making. The final bump foil geometry details are provided in Figure 2 and its calculated stiffness is given in Table 1. The first step in the experimentation is to measure the nominal/initial clearance of the bump foil bearings manufactured. To measure the clearance and lift-off speed, a dedicated test rig is designed and fabricated. The schematic of the test rig is shown in Figure 3. It consists of an overhung shaft supported on two highspeed rolling element bearings at the center. On one end of the shaft, air turbines drive is mounted to spin the shaft at desired speed by passing the compressed air over the turbines. On the other end of the shaft, a sleeve of 10mm inner diameter is introduced whose outer diameter can be varied depending on the test bearing inner diameter. This makes the test rig versatile to accommodate bearings of different sizes with minimal modification"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002163_s40430-020-02796-3-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002163_s40430-020-02796-3-Figure4-1.png",
        "caption": "Fig. 4 Installation diagram of bearing ring block",
        "texts": [
            "\u00a0(22) using successive over-relaxation (SOR) Iteration Method: where \u03b2 is super relaxation factor, 1 < \u03b2 < 2; k is number of pressure iterations. When pressure is iterated to a certain number of times, program stops iterating when formula (25) is satisfied \u03b6 is pressure convergence accuracy, is 1.0 \u00d7 10\u201310. Diagram of simulation calculation process is shown in Fig.\u00a01. The picture of test setup is shown in Fig.\u00a02, bearing ring block is shown in Fig.\u00a03, and installation diagram of bearing ring block is shown in Fig.\u00a04. Friction torque of bearing ring block can be obtained by torque equipment as shown in Fig.\u00a02. During tests, load P (specific pressure, Pa) is set to 0.2\u00a0MPa by weight Fblock (unit: N). (22) B1i,jpi+1,j + B2i,jpi\u22121,j + B3i,jpi,j+1 + B4i,jpi,j\u22121 \u2212 B5i,jpi+1,j = B6i,j (23) \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 B1i,j = \ufffd h 3 \ufffd i+0.5,j B2i,j = \ufffd h 3 \ufffd i\u22120.5,j B3i,j = \ufffd d l \ufffd2\ufffd \u0394 \u0394 \ufffd2 \ufffd h 3 \ufffd i,j+0.5 B4i,j = \ufffd d l \ufffd2\ufffd \u0394 \u0394 \ufffd2 \ufffd h 3 \ufffd i,j\u22120.5 B5i,j = B1i,j + B2i,j + B3i,j + B4i,j B6i,j = 3\u0394 \ufffd\ufffd h \ufffd i+0.5,j \u2212 \ufffd h \ufffd i\u22120.5,j \ufffd (24) p k i,j = ( B1i,jp k\u22121 i+1,j + B2i,jp k\u22121 i\u22121,j + B3i,jp k\u22121 i,j+1 + B4i,jp k\u22121 i,j\u22121 \u2212 B6i,j B5i,j ) + p k\u22121 i,j (25) \u2211n j=2 \u2211m i=2 \ufffd\ufffd\ufffdpki,j \u2212 pk\u22121 i,j \ufffd\ufffd\ufffd\u2211n j=2 \u2211m i=2 \ufffd\ufffd\ufffdpki,j \ufffd\ufffd\ufffd \u2264 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:56 1 3 Page 5 of 10 56 where P is specific pressure of bearing ring block, Pa; S is area of bearing ring block, m2; R2 is outer diameter of bearing ring block, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000687_s10846-020-01149-5-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000687_s10846-020-01149-5-Figure1-1.png",
        "caption": "Fig. 1 An example of boundary singularity of the anthropomorphic robotic arm",
        "texts": [
            " In the experiments, we design a trajectory that pass through non-singular region, boundary singularity, and internal singularity, and examine the ability of the proposed algorithm for real-time control. Firstly, the boundary singularity handling simulation is demonstrated below. The desired trajectory is designed to travel from a regular configuration to the workspace boundary, and then back to the non-singular region (threshold is set to s0 = 10 \u22124). We consider a singular configuration q\u00bc 0 \u2212 \u03c0 7 0 \u2212 \u03c0 2 0 0 T at the boundary singularity (Fig. 1). In this configuration, rank(J) = 5 and the rotated Jacobian matrix in the frame 0 can be shown as Js\u00bcRT s J q\u00f0 \u00de\u00bc 0 \u22121:2 0 0:8 0 0:5 0:5 0 0 0 0 0 0 0 0 0 0 0 \u22120:4 0 0 0 0 0 0 \u22121 0 1 0 1 0:9 0 1 0 1 0 2 6666664 3 7777775 \u00f024\u00de We now observe that the singular direction of (24) is obviously in the zs-direction due to the third row of (24) is all zero. The control process of the end-effector of this simulation is shown in Fig. 2, where the sagittal apex of the blue line represents the desired trajectory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002309_j.ijepes.2021.106860-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002309_j.ijepes.2021.106860-Figure7-1.png",
        "caption": "Fig. 7. Cross-section of auxiliary PMG pole with necessary parameters for calculation.",
        "texts": [
            " \u03d5\u03c3 = VsK (6) Total magnetic conductivity K is defined as a sum of the magnetic conductivity between the pole end plates (Kp), and the magnetic conductivity between the magnets Km, Kp = [ Lplp 0, 8ap + 2lplog ( 1 + \u03c0 2 bp ap )] 10\u2212 6 (7) Km = [ Lmlm 1, 6am + lmlog ( 1 + \u03c0 2 bm am )] 10\u2212 6 (8) F. Zec et al. International Journal of Electrical Power and Energy Systems 129 (2021) 106860 where Lp represents the length of the pole end plate, and Lm represents the length of the pole magnets [16]. All other parameters for calculation are shown in Fig. 7. Stator magnetic voltage is defined as: Vs \u2245 2H\u03b4\u03b4+ 2HZhZ . (9) Calculated value of stray flux \u03d5\u03c3 is 0.00157 Vs (Eq. (6)), represents about 5% of the stator flux \u03d5sf (0.0297 Vs), which was calculated using the induced voltage formula: where U represents the PMG no load voltage on the stator terminals (449\u0305\u0305 3 \u221a V). Because the value of stray flux can be neglected, additional ampereturns are not required. Next step is the calculation of the required magnetization current, based on the calculated value of ampere-turns and the known winding connection of PMG"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002359_jestpe.2021.3059280-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002359_jestpe.2021.3059280-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of a three-sector three-phase PMSM.",
        "texts": [
            " Then, the comprehensive suppression method for PWM noise is illustrated in Section III. In Section IV, the PWM elimination principle of PWM harmonics using the proposed method is elaborated, and the advantages of the drive system are discussed. Afterwards, experiments are conducted on a two-sector three-phase PMSM controlled by a digital signal processor control platform in Section V. Finally, conclusions are drawn in Section VI. II. INTRODUCTION OF THE PROPOSED DRIVE SYSTEM A typical multi-sector three-phase PMSM is exhibited in Fig. 2 and can be divided into three sectors(6 slots and 2 poles per sector) [3]. The independent sectors lead to lower end turns of the stator winding, making the phase resistance of the motor smaller. Besides, mutual inductances among different three-phase winding sets become extremely low and negligible due to winding displacement[1]. Thus, the magnetic decoupling between sectors allows each submotor to operate independently. Consequently, a fault of one subunit introduces only a fraction of torque disturbance and has almost no influence on other subunits, except that every other subunit must produce k/(k-1) current to fulfil the rated torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001770_dvm49764.2020.9243883-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001770_dvm49764.2020.9243883-Figure8-1.png",
        "caption": "Fig. 8. The topology optimized Valve Island",
        "texts": [
            " Combination of topology optimization and FEA analysis in Ansys Additive, Simufact Additive allow optimizing a part design and SLM technological regimes (laser power, scan speed, hatch distance) to eliminate form deviation caused by Authorized licensed use limited to: Rutgers University. Downloaded on May 18,2021 at 15:37:54 UTC from IEEE Xplore. Restrictions apply. 978-1-7281-7526-3/20/$31.00 \u00a92020 IEEE temperatures gradient and internal stress. Additionally, reduce the operational cost and lead production time. Samara University Case Study We developed the technology of redesign by topological optimization and SLM methods for manufacturing a valve island (Fig.8). The volume of the optimized design part was 30368 mm3, which is 30% of the initial volume. The initial mass of the structure was also reduced by 70% to 0.2383 kg. The Software Simufact Additive has been used for finite element simulation of part form deviation caused by thermal strain, which are a result of temperatures gradient during SLM process (Fig.9). An analysis of the results showed that the greatest total deformations were 0.21 mm on the end surface of the part. The total deformation of the central holes was 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001034_s40815-020-00869-y-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001034_s40815-020-00869-y-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the bicycle model",
        "texts": [
            " lf and lr are the distances from front and rear (FR) wheel axles to the center of gravity (CG), respectively. Iz stands for moment of inertia. vx and vy represent the vehicle longitudinal velocity and lateral velocity, respectively, where vx is a constant. The yaw moment Mz is written as [45] Mz \u00bc \u00f0Fxflcosdf \u00fe Fxrl\u00dels \u00f0Fxfrcosdf \u00fe Fxrr\u00dels where df and Fxi denote the steering input angle and the longitudinal force, respectively. Since the four wheels of electric vehicle model are symmetrical on both sides, it is reasonable to give the bicycle model depicted in Fig. 2 for the simplification of controller design. Then, we construct the following vehicle handling dynamic model [25]: mvx _b\u00f0t\u00de \u00bc Fyf \u00f0t\u00de \u00fe Fyr\u00f0t\u00de mvxXz\u00f0t\u00de Iz _Xz\u00f0t\u00de \u00bc lf Fyf \u00f0t\u00de lrFyr\u00f0t\u00de \u00feMz\u00f0t\u00de ( \u00f01\u00de where the FR lateral forces are represented by Fyf and Fyr, respectively, and they are given by Fyi\u00f0t\u00de \u00bc fiai\u00f0t\u00de; \u00f0i \u00bc f ; r\u00de \u00f02\u00de where ff and fr represent the cornering stiffness of FR wheel, respectively. In addition, the FR slip angles can be denoted by af and ar, respectively, and they satisfy af \u00f0t\u00de \u00bc df \u00f0t\u00de lfXz\u00f0t\u00de vx b\u00f0t\u00de ar\u00f0t\u00de \u00bc lrXz\u00f0t\u00de vx b\u00f0t\u00de: \u00f03\u00de By substituting (2) and (3) into (1), the state-space form of vehicle handling dynamic model (1) is deduced as _v\u00f0t\u00de \u00bc Av\u00f0t\u00de \u00fe B1x\u00f0t\u00de \u00fe B2u\u00f0t\u00de where A \u00bc ff \u00fe fr mvx 1 lf ff lrfr mv2x lf ff lrfr Iz l2f ff \u00fe l2r fr Izvx 2 6664 3 7775 B1 \u00bc ff mvx lf ff Iz 2 664 3 775; B2 \u00bc 0 1 Iz 2 4 3 5; x\u00f0t\u00de \u00bc df \u00f0t\u00de v\u00f0t\u00de \u00bc b\u00f0t\u00deXz\u00f0t\u00de\u00bd T ; u\u00f0t\u00de \u00bc Mz\u00f0t\u00de: The following IT2 fuzzy model depicts the electric vehicle systems with uncertainties"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002118_s00521-020-05554-7-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002118_s00521-020-05554-7-Figure1-1.png",
        "caption": "Fig. 1 Lower extremity exoskeleton robot design",
        "texts": [
            " HAL-5 will help good wearers in terms of results while standing up, walking, climbing stairs and a number of other daily living activities [62]. For these reasons, 85% of the designed LLEs are electrical actuators, 4% are hydraulic actuators, and 1% are pneumatic actuators [63]. Electric motors are preferred as the drive systems because of the easy application of advanced control techniques to achieve the best performance [64, 65]. Within this scope, in this study, the design of the lower extremity exoskeleton robot is shown in Fig. 1. The lower- limb strengthening robot\u2019s mechanical design was developed using the SolidWorks software. For mechanical design, the range of motion required was taken from [66]. The hip joint will move between - 30 and 30 . Like the hip joint, the knee joint will move between - 5 and 60 during the gait pattern [67]. The experimental setup\u2019s total weight is 8 kg. The weight of all the lower extremity robot systems (building, motors, batteries, etc.) is 20 kg. It was necessary to apply approximately 200 Nm of torque to the joints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure4-1.png",
        "caption": "Fig. 4. A detailed view of the mesh quality in the 2-D model of Permanent Magnet Synchronous Machine.",
        "texts": [
            " The geometrical of 2D and 3D model are depicted in Fig. 2 and Fig. 3. 682 Authorized licensed use limited to: Raytheon Technologies. Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. During the FE simulation of permanent magnet synchronous machine, the meshing is essential process which is done to descretize the geometry developed into small number of parts called cells. A quarter geometry of the PMSM is subdivided into triangular elements (1000 elements). As obvious from Fig. 4 and Fig. 5 both models give relatively the detailed view of the mesh quality in the 2D and 3D model of the PMSM. Fig. 6 shows the electromagnetic torque calculated with the 2D and 3D FE models: The torque obtained from this model is also shown in Fig. 6. It can be seen that the shape of the electromagnetic torque differs noticeably from the others. The blue curve (3D) obviously lies much closer to the measurement. It is because the machine is relatively short and the edge effects are (not considered in 2D) not negligible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000771_iccia49288.2019.9030988-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000771_iccia49288.2019.9030988-Figure6-1.png",
        "caption": "FIGURE 6. Power test bench of the Gycom gyrotron. The gyrotron output power is the sum of the power of dummy load, bellows, waveguides, waveguide switch, MOU, miter bend, and output window. Two power monitor miter bends are used. One power monitor miter is situated between MOU and dummy load, the other one which is similar to the one used for CPI gyrotron is situated between the first power monitor bend and tokamak. The gyrotron output power can be got by calorimetric method.",
        "texts": [
            " And the coupling of the power monitor miter bend for Gycom gyrotron is not clear, but it will not be the difficulty for us to realize the real time power monitoring as shown below. VOLUME 4, 2016 2169-3536 2016 IEEE. Translations and content mining are permitted for academic research only. Personal use is also permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. 5809 A. INCIDENT WAVE POWER MONITORING The 3D pictures of the power test benches for Gycom gyrotron and CPI gyrotron are shown in Fig.6 and Fig.8. One power monitor is situated between the MOU and the dummy load for CPI gyrotron. This power monitor miter bend is situated before the polarisers and it picks up only 5810 VOLUME 4, 2016 one polarisation. The output wave from the MOU is H-plane linearly polarised. As shown in Fig.6, two powermonitors are used for Gycom gyrotron. One is situated between the MOU and the dummy load, the other one which is similar to the one used for CPI gyrotron is situated between the first power monitor miter bend and tokamak. Both two powermonitors pick up only one polarisation, and both are situated before the polarisers. The output wave from the MOU is E-plane linearly polarised for No.1 power monitor miter bend, and thus the wave is H-plane linearly polarised for No.2 power monitor miter bend",
            " The exponential function can represent the relationship between the RF output power and the detector output voltage to some extent. But the dispersion is large which may due to the characteristic of sensitivity of detector being vulnerable to the temperature. And actually, the varying of mode purity at the output of the MOU could cause such a scatter. For confirming whether the power monitor or the gyrotron is to blame for the scatter, the other power monitor which is similar to the one used for CPI gyrotron is used for Gycom gyrotron. As shown in Fig.6, it is situated between the first power monitor miter bend and tokamak. Firstly, the pulse is injected to dummy load, so we can get the output power by calorimetric method. Then we keep all the same gyrotron operation parameters but inject the pulse to plasma, 5814 VOLUME 4, 2016 got the output voltage of the No.2 power monitor for Gycom gyrotron. The test results are shown in Fig.18. The RF output power can be expressed by a linear function, Pout \u2248 17.29+ 236.49 \u2217 Vdetctor (16) The Adj. R-Square is about 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000450_sensors43011.2019.8956497-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000450_sensors43011.2019.8956497-Figure1-1.png",
        "caption": "Fig. 1. (a) The GO drop casted on to the SPCE (b) ERGO-SPCE with a 3D printed well attached for precise sample loading",
        "texts": [
            "6) and SA powder were purchased from Millipore-Sigma, MO, USA. Screen printed carbon electrodes (Zensor) were purchased from CHI instruments, TX, USA. DI water with resistivity of 18.2 M\u2126cm was used for making solutions unless otherwise specified. Salicylate liquiUV test kit was purachsed from Stanbio labs, TX, USA. CHI660E electrochemical workstation was used for DPV and UNICO SQ2800 UV/VIS spectro-photometer was used for measuring reference SA level in orange juice with the Salicylate liqui-UV test kit. Fig. 1 shows the proposed ERGO-SPCE sensor. The sensor fabrication procedure involved, first, modifying the electrode with GO nano-sheets. GO flakes were dispersed in deionized water (DI water) and thoroughly sonicated for 2 hours at room temperature to obtain a suspended solution of GO with a concentration of 1 mg/ml. The SPCE surface was cleaned with ethanol and rinsed with DI water followed by plasma cleaning using air plasma for 30 seconds at 15 W power. 7 \u00b5l of GO solution was then drop-casted on to the SPCE and was left to air dry at room temperature for 24 hours"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure27-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure27-1.png",
        "caption": "Fig. 27. Axial load-displacement test bench.",
        "texts": [
            " For design 2 (25\u00b0 contact angle 4PBB) and design 5 (35\u00b0 contact angle 4PBB), the load-displacement curve under 3tF is shown in Fig. 26. For 4PBB with two different contact angles, the displacement at 1000 N drops from 104 \u03bcm (25\u00b0 contact angle) to 81 \u03bcm (35\u00b0 contact angle). And with the increase of load, the loaddisplacement curve of large contact angle bearings tends to be more horizontal (increased rigidity). It can be seen that the increase of contact angle can reduce the displacement under load to a certain extent. The bearing axial load-displacement test bench is shown in Fig. 27. The bearing inner ring and the shaft are rigidly fixed, and the outer ring is rigidly connected to the housing, as shown in Fig. 28. Axial load is applied to the shaft in both directions, and the axial displacement of the shaft is measured by a displacement sensor. The measurement software records the data from the load cell and the displacement sensor at the same time to obtain the axial load-displacement curve. The bearing tilt load-displacement test bench is shown in Fig. 29. The schematic figure of measurement is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001612_13621718.2020.1825060-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001612_13621718.2020.1825060-Figure1-1.png",
        "caption": "Figure 1. Schematic sketch of physical model.",
        "texts": [
            " In this work, the average amplitude and pitch of surface ripples after solidification are more realistically scaled by further accounting for solidification rate and oscillations of fluid flow. A thermal fluid model is also proposed and solved by using COMSOL code. Successful comparisons between the scaled and predicted ripples are required to confirm and control surface quality of welding, additively manufacturing and nanotechnology. In this model, a workpiece moves with a constant speed U in the positive x-direction, as illustrated in Figure 1. The physical model composed of air, liquid and solid is also applicable for a stationarymetal subject to incident energy flux moving at a scanning speed U in the negative x-direction. Without loss of generality, the major assumptions made are the following: (1) The model system with incompressible air is twodimensional in a steady periodic state in a moving coordinate of \u03be [24]. The assumption is often used as a first approximation to relevantly reveal basic mechanisms of a dynamic process. (2) The incident flux is modelled by introducing convection heat transfer on the left and top boundaries of the computational domain (see Figure 1). Initial temperatures in air, liquid and solid were thus chosen to be 1000, 950 and 900K, respectively. (3) The initial molten pool is a thin layer in a rectangle shape within solid. The simplified pool shape was chosen in order to save computer time to produce ripples. Even though the shape seems different from a realweld pool, the analysis is focused on the corner region, which is quite similar to the realistic welding process. (4) Solid speed and oscillation velocity in the liquid layer at the left boundary can be specified as u = \u23a7\u23a8 \u23a9 U + f (u \u2212 U) \u221a\u2223\u2223 y 0",
            " The fluid flow and heat transfer modules in COMSOL code with boundary conditions listed in Table 1 were used to solve Equations (2)\u2013(4), and the separated multiphase flow models to solve Equation (7). Grid points were 3.6 \u00d7 105. Relative and absolute errors were, respectively, 0.001 and 0.05. To ensure convergence of Equation (7), variables in\u03c7 < 0.01 sm/kg, and interface thickness parameter \u03b5 = 2.5 \u00d7 10\u22125 m. Time and spatial derivatives in Equations (2)\u2013(4) in a moving coordinate \u03be , y, t, where \u03be = x \u2212 Ut (see Figure 1), yield \u2202 \u2202t \u2223\u2223\u2223\u2223 x = \u2202 \u2202t \u2223\u2223\u2223\u2223 \u03be \u2212 U \u2202 \u2202\u03be , \u2202 \u2202x = \u2202 \u2202\u03be (8) where the term \u2202/\u2202t|\u03be = i\u03c9 is for a steady periodic system [24]. The Bernoulli equation by integrating from Equation (3) over coordinate \u03be along the free surface between the left boundary to edge of the pool then leads to \u03c1u2s \u2212 \u03c1Uus \u2212 (\u03c1u2e \u2212 \u03c1Uue) \u2248 pe \u2212 ps + Kloss\u03c1u2s (9) where the last term on the right-hand side represents mechanical energy losses due to viscous dissipation, steady periodic effects, and other irreversible processes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure5-1.png",
        "caption": "Fig. 5. Demagnetization ratio distribution for STPSPM machines without eccentricities at 16ms during 3PSC. (a) 10-pole/12-slot. (b) 10-pole/9-slot.",
        "texts": [
            " The demagnetization area and level can be indicated by demagnetization ratio distribution, which is defined as 0 0 r r demag r B B r B \u2212 = (1) where Br0 is the PM inherent remanence and Br is the remanence after demagnetization. In addition, post-demagnetization machine performances can also be calculated. Therefore, after 3PSC ends, iq=8A is again applied for a mechanical period to obtain the post-demagnetization UMF. The 3PSC current waveforms are shown in Fig. 4 and demagnetization ratio distribution is shown in Fig. 5. It can be seen that the demagnetization distribution is rotationally symmetrical for the 10-pole/12-slot STPSPM machine. Hence, there is no post-demagnetization UMF, as shown in Fig. 6. For the 10-pole/9-slot STPSPM machine, the demagnetization distribution is rotationally asymmetric, as shown in Fig. 5(b). Therefore, as shown in Fig. 7, the additional 1st order harmonic occurs in Fx and Fy, where Fx and Fy are the x and y components of UMF, respectively. It can be seen the effect of demagnetization on UMF for asymmetric STPSPM machines is similar with the influence of rotating eccentricity. In addition, since the UMF magnitude is calculated as 2 2 x yF F F= + , (2) the constant component of the UMF magnitude is increased and the 9th order harmonic is added, as shown in Fig. 7 (e) and (f). Authorized licensed use limited to: California State University Fresno",
            " (a) (b) -40 -20 0 20 40 0 4 8 12 16 20 24 C u rr e n t (A ) Time (ms) C u rr e n t (A ) A B C Static eccentricity No eccentricity -40 -20 0 20 40 0 4 8 12 16 20 24 C u rr e n t (A ) Time (ms) Rotating eccentricity No eccentricity B C A Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. The demagnetization distributions due to 3PSC are shown in Fig. 13. For symmetrical SPM machines, when there is no eccentricity, the demagnetization distribution is rotationally symmetrical, as shown in Fig. 5(a). However, with static or rotating eccentricity, the demagnetization is not rotationally symmetrical and gathers around the PMs facing larger air-gap length. This can be explained by open-circuit and armature field working points, and the influence of static eccentricity on symmetrical STPSPM machine is analyzed as an example. It is defined that positive x direction in the stator reference is 0 mechanical degrees whilst the counter-clockwise direction is the positive direction, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure15-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure15-1.png",
        "caption": "Fig. 15. Temperature distribution with axial water jacket cooling for J= 15.6 A(rms)/mm2",
        "texts": [
            " Since only one channel carries the coolant, the entire flow volume is through this channel, 1087 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 19,2021 at 14:42:00 UTC from IEEE Xplore. Restrictions apply. thus keeping the total flow rate the same as discussed in Section III. As presented in Fig. 14, the maximum temperature is well above the permissible limits. Thus, the maximum allowable slot current density using axial water jacket cooling for the same temperature rise of 80 \u25e6C, as seen in Fig. 15, is 15.6 A/mm2. This current density corresponds to 50% lower compared to that of the WELC method. This paper proposes a low thermal resistance winding embedded liquid cooling (WELC) concept for slotless motors leveraging the space within the non-magnetic winding support for efficient heat extraction out of the windings, and thereby achieving higher current densities. The developed thermal management method for slotless motor enables a lightweight design giving very high power density at the system level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure33-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure33-1.png",
        "caption": "Fig. 33. Load-displacement curve of the solid block.",
        "texts": [
            " To measure the displacement of the bench system under the measuring load, a solid block (shown in Fig. 31) is specially made with the same dimensions of the test bearing (6004 inner diameter, outer diameter, and width are 20 x 42 x 12 mm). The deformation of the solid block under the axial test load condition can be calculated using the FEM method, as shown in Fig. 32. The maximum displacement of a solid block under the load of 1000 N is 0.4 \u03bcm. The measured axial load-displacement curve of the solid block is shown in Fig. 33. According to Fig. 33, the axial displacement of the solid block under the axial load of \u00b11000 N is -3.6 \u03bcm and 3.7 \u03bcm. Considering that the theoretical deformation displacement of the solid block is only 0.4 \u03bcm, it can be considered that the dis- placement under the load at this time all comes from the deformation of the fixture. The measured bearing axial loaddisplacement curve will be corrected in \u00b11000 N by deducting the curve in Fig. 33. Note that the load and unload curve in Fig. 33 does not overlap. The reason is that the internal friction of the device itself causes the displacement to lag when returning. When the solid block is subjected to a loading force of \u00b1100 N at the position of the loading arm as shown in Fig. 29, the deformation is shown in Fig. 34 using FEM analysis. According to the analysis results in Fig. 34, the maximum radial displacement of the inner bore surface of the bearing is 0.24 \u03bcm and -0.21 \u03bcm, so a radial displacement of 5.6 \u03bcm will occur at the point of the measuring arm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000848_5.0000820-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000848_5.0000820-Figure7-1.png",
        "caption": "FIGURE 7. Location of modified part in the frame",
        "texts": [
            " Figure 6 shows the value of intrusion distance for the existing material. According to result in the Figure 5 and Figure 6, the frame must be modified to enhance its performance since the structure can be considered as fail. The value of intrusion and also the stress were not fulfilling the requirement of ECE 66 standard. Some part of the frame has been modified by increasing its thickness to reduce the stress and also the value of intrusion. List of modification was tabulated in Table 3. The location of modified part in the frame were depicted on Figure 7. 030153-5 The simulation was done for the modified frame and the result for displacement, stress and also intrusion were depicted in Figure 8, Figure 9 and Figure 10 respectively. Based on the result of simulation, the displacement of frame was decrease. The magnitude of maximum deflection is 359.4 mm, and minimum displacement is 39.94 mm. 030153-6 The von Misses stress actually increase compared to the existing frame, but during the rolling accident, the collapse part is needed to absorb the impact energy from the external load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002684_j.engfailanal.2021.105424-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002684_j.engfailanal.2021.105424-Figure2-1.png",
        "caption": "Fig. 2. A pair of fracture surfaces from the failed screw shaft.",
        "texts": [
            " The tensile test was carried out on a ZWICK100KN with gauge dimensions of 18 mm in length, 6 mm in width and 2 mm in thickness. A constant strain rate of 10-3 s\u2212 1 was used, and at least four samples were measured to guarantee reproducibility and statistical relevance. The diagram of the tensile samples is shown in Fig. 1. The shaft studied in the present study had been in service for ten years before failure, and the fracture occurred at the root of the failed screw, that is, the reduced diameter of the screw shaft. The overview of the fractured shaft is shown in Fig. 2. Fig. 3(a) shows the visual inspection of the failed screw shaft. It can be observed that the outer surface of the side near the fracture surface was seriously worn, and clear furrows can be observed. Fig. 3(b-d) displays the enlarged images of the failed screw shaft along with different parts of the circumferent surface. In addition to deep furrows (yellow arrow), in some places, the furrows were first worn in and then smoothed out (red arrow), and some pits can also be observed (blue arrow). These visual observations show that the screw shaft suffers from serious wear during the operating process, which has a great impact on the service life of the shaft, especially on its M"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001825_15397734.2020.1841003-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001825_15397734.2020.1841003-Figure3-1.png",
        "caption": "Figure 3. The hybrid composite shaft power turbine rotor-bearing system.",
        "texts": [
            " The critical speed analysis is carried out using ANSYS SHELL281 element. The details of the hollow steel shaft with both the ends supported on the bearing supports as per Genta and Amati (2010) are given in Table 2. The material loss factor is modeled as proportional stiffness damping. The analysis is carried out using MATLAB finite element code and verified by ANSYS using BEAM188 element. The geometric details of power turbine rotor-bearing system under study are given in Table 3. The finite element model is also shown in Figure 3. The numerical rotor dynamic analysis methodology followed in the present work is explained in the following sections. Initially, a parametric analysis is performed to shortlist the best material configuration for the application followed by critical speed analysis, stability analysis, and unbalance response analysis. ANSYS SHELL281 element is used to carry out the parametric analysis where the modal and torsional stress analysis are performed to select the best shaft configuration by parametrically varying the number of layers, layer sequences, and layer thickness for the composite materials given in Table 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002037_j.engfailanal.2020.105126-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002037_j.engfailanal.2020.105126-Figure18-1.png",
        "caption": "Fig. 18. Lower arm of the structure (arm 3).",
        "texts": [
            " If the input is taken into account in the Table 6, then, the value of X, is the thickness of ice X for a coefficient of variation of 20% and a probability of 0.2% to be exceeded in a year or a return period of 500 years using the following Eq. (8) and Table 7. X = X\u2019 \u2212 0.45\u03c3 + \u03c3 1.282 [ \u2212 ln( \u2212 ln(1 \u2212 p(X)) ) ] (8) The designs were made taking into account stress, dead and wind loads for suspension structures (Fig. 8 and Fig. 9) are the Fig. 14, Fig. 15, Fig. 16 and the retention (Fig. 11 and Fig. 12) structures are in the Fig. 17, Fig. 18. The wind load was estimated following the procedure established by this paper in the section 2 and Eq. (9). Load = 0.613 \u00d7 V2Vkz \u00d7 GRF \u00d7 I \u00d7 Dd \u00d7 A (9) Where: Fig. 11. Retention structure for normal condition. Fig. 12. Retention structure for construction condition. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 0.613: It is the constant for metric unit for the density in the air. It reflects the air density for 1 atmosphere, 15 C, with mercury pressure 760 mm, in the Table 8",
            " 27 and Fig. 28. In the Fig. 14, the suspension tower is modeled with finite element; each part of the tower has been built for the validation of the loads according to the section 3. In the Fig. 15, the material for retention tower has built, with the material influence for the calculations and constraints as elastic module, yield stress, steel type and components. Each component of the tower has been described in the Fig. 16 with the tower top. Fig. 17 for both upper arms, it applies to arm 1 and 2, Fig. 18 with the lowest arm of the tower, Fig. 19 in the upper body between arm 1 and 2, Fig. 20 for the body structure between arm 2 and 3, Fig. 21 and Fig. 22. Fig. 14. Tower model in finite element. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 The stage division of the ice disaster is composed of four stages: Preconditions (associated to the pre-failure period), steady-state progression and multiple failures (during disaster period) and restoration (post-failure period) [38]; recent papers have indicated effective resilience enhancement framework for towers installed, however, the recommendations for new assets are not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002781_s00366-021-01416-5-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002781_s00366-021-01416-5-Figure13-1.png",
        "caption": "Fig. 13 Push-pulling the motorcycle triple clamp model and the modeling results",
        "texts": [
            " Besides, by comparing the modeling results of Siemens NX and the proposed method, the proposed GTI detection method is seen to be effective. Case studies 1\u20133 focused primarily on the GTI detection module, and thus the effectiveness of the presented method as a whole was not sufficiently demonstrated. For this reason, three more case studies that are comprehensive were carried out. Considering that complexity of push-pulling a model lies in how many critical points it needs to cross, and how complex the associated GTI configurations are, case study 4 (Fig.\u00a013) analyzed a situation with multiple critical points. Push-pulling the triple clamp model in Fig.\u00a013 involved 10 critical points in total, and some of them occurred concurrently. As a result, the GTI configurations are very complex, and the associated detection task is challenging. Siemens NX was only able to successfully cross the first two critical points, and SpaceClaim was not even able to cross a single critical point (yielding a weird resulting geometry), while the proposed method can correctly detect all the critical points and resolve the generated GTI. In case study 5 (Fig.\u00a014), instead of linearly adding more critical points to the push\u2013pull move, the comprehensiveness was attained by push-puling a same model under various situations: (1) push\u2013pull the blue faces and stop in between the first and second critical pointsz; (2) push\u2013pull the blue faces, and stop in between, then continue the push\u2013pull until the end; and (3) push\u2013pull the blue faces until the end"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001885_icem49940.2020.9270870-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001885_icem49940.2020.9270870-Figure3-1.png",
        "caption": "Fig. 3. Perspective view of prototype linear compressor.",
        "texts": [
            " 2 shows a cross-sectional view taken on a plane along the A-B-C-D line in Fig. 1. The test linear motor is a single-phase machine, and the armature has a structure in which a mover having permanent magnets is sandwiched vertically. This structure, in which the armatures are arranged facing each other, has a feature in that the magnetic attractive force is canceled [12]. Moreover, since each magnetic pole has an independent structure, the magnetic pole length and the pole pitch can be easily changed. Fig. 3 shows the configuration of a test linear compressor having two armatures, and Table I shows the specifications. Assist coil springs, i.e., resonant springs, are connected to both sides of the mover, and the mover has a structure in which the mover reciprocates the piston by the electromagnetic exciting force of the linear motor. A piston connected to the mover draws refrigerant or air into a cylinder and performs compression and discharge. The prototype linear compressor has two armatures, shown in Position Sensor-less Resonant Frequency Estimation Method for Linear Compressor with Assist Springs Takahiro Suzuki, Masaki Koyama, Shuhei Nagata, Wataru Hatsuse, Masatsugu Takemoto, and Satoshi Ogasawara A Authorized licensed use limited to: Auckland University of Technology"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002435_s40799-021-00451-7-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002435_s40799-021-00451-7-Figure2-1.png",
        "caption": "Fig. 2 Schematic of weld bead for calculating bead coefficient",
        "texts": [
            " Three possible cases of bead geometry are considered: No penetration, under penetration and full penetration and their corresponding areas determined are shown in Table 5. There are three factors namely angles of the bevels, length of the weld root and constant a (value adjusted according to the weld depth) that determine volume of the weld. In CMT process, the material deposition is integrated with the filler wire feed motion. This advanced feature allows the transfer of molten metal in droplets rather than spattering. The controlled molten material droplet detachment in CMT results in the circular-shaped cross-sectional beads (Fig. 2). The bead coefficient (BC) is calculated from the total volume deposited (Vd) and the nominal value of deposition (V). It is expressed as the follows. BC \u00bc Vd\u2212Vj j V \u03b1 \u00bc Sd\u2212Sj j S \u03b1 \u03b1 \u00bc S2 S1 \u00fe S3 \u00f05\u00de where \u03b1 is the penalising factor, Sd is the crosssectional area of total deposited material, S is the nominal deposition and S1, S2and S3are the sectioned crosssectional areas as shown in Fig. 1. The factor \u03b1 in Eq. (5) penalizes the inadequate penetration, as penetration is the major influencing factor of weld strength"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000284_j.compeleceng.2019.106534-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000284_j.compeleceng.2019.106534-Figure4-1.png",
        "caption": "Fig. 4. Mesh generation and boundary setup of the BIM.",
        "texts": [
            " (3) The field quantities are assumed periodically varying and the temperature effect on the material is neglected. Based on these assumptions, the 2-D transient field with magnetic vector potential A z can be described as follows [14, 15] . { \u2202 \u2202x ( 1 \u03bc \u2202 A z \u2202x ) + \u2202 \u2202y ( 1 \u03bc \u2202 A z \u2202y ) = \u03c3 \u2202 A z \u2202t \u2212 J s A z | 1 , 2 = 0 (1) where \u03bc is the medium permeability, \u03c3 is the conductivity, J s is the source current density, 1 is the stator outer circumference boundary, and 2 is the rotor inner circumference boundary. The mesh generation and boundary setup of the BIM is shown in Fig. 4 . Eq. (1) can be transformed into a ubiquitous equation. After the regional unit is divided and linearly interpolated, a nonlinear equation system can be obtained as KA = P (2) where K is the coefficient matrix, A is the node magnetic vector potential, and P is the excitation. Then, by solving the equations, the vector magnetic potential of each node is gained. Finally, the electromagnetic quantities in each region of the motor can be calculated. Regarding the external circuit model, the induced electromotive force of stator winding is expressed with magnetic vector and the voltage balance equation of stator winding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002726_j.ymssp.2021.107837-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002726_j.ymssp.2021.107837-Figure3-1.png",
        "caption": "Fig. 3. The SEA testbed.",
        "texts": [
            " This guarantees that all relevant variables above are semiglobally bounded for all t P 0. In addition, from (51), it leads to 0 6 V 6 vmax f \u00fe V\u00f00\u00de vmax f e ft \u00f052\u00de Then, from (52), it further leads to lim t!1 V \u00bc vmax f . Therefore, the tracking error s1 can eventually converge to an arbitrarily small neighborhood around the origin by increasing f. This completes the proof. The proposed control strategy should be tested in real experiments before being applied to practical applications. The SEA testbed on which the experiments are conducted is shown in Fig. 3. It mainly consists of a brushless AC motor, a harmonic reducer, a self-made torsional spring, and a load bar. Two incremental encoders, with a resolution of 250000pulses=rev and 1048576pulses=rev, are mounted to measure the angular positions of the motor axis and the load axis, respectively. A personal computer (PC) on which the real-time operating system (RTOS) runs is employed to execute the control algorithms. The National Instruments PCIe-6321 card installed on the PC receives the two quadrature encoder signals and sends the analog control command to the motor drive at 1kHz control frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001845_s00202-020-01132-1-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001845_s00202-020-01132-1-Figure2-1.png",
        "caption": "Fig. 2 Leakage of airflow in the axial orientation [4]",
        "texts": [
            " The TEFC machine is cooled by passing the coolant over spacing among the semiopen fins of the housing. In comparison with some other active cooling systems, such as the liquid cooling systems, in which the coolant flow rate is predicted with reasonable accuracy, the estimation of the airflowvelocity from the housing of a TEFC machine is a challenging task [6]. There are several reasons to describe this statement. The distribution of the airflow produced by the fan is not uniform in every fin channel [6, 7]. According to Fig. 2, by leaking the airflow from semi-open fin channels, the air speed starts to reduce in the axial orientation [6]. Moreover, some of the fin channels of TEFC housing are blocked by the terminal box and bolt lugs, which terminates in the unbalanced distribution of the airflow about the machine housing [7, 8]. Therefore, calculating the heat transfer coefficient from the housing of TEFC is the conjugate problem, which consists of fluid flow and heat flow problems [6]. Different studies consider the thermal analysis of TEFC machines"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003238_s12206-021-0829-0-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003238_s12206-021-0829-0-Figure2-1.png",
        "caption": "Fig. 2. Planetary gear opening.",
        "texts": [
            " The ring gear is a semi-floating member that is connected to the output shaft by bolts. The torque transmission bracket bears the gravity of the gearbox, and torque is generated when the gears mesh. To improve the oil-return capability of the fan-driven gearbox, the influence law of the planetary gear hole parameters on the heat transfer performance of the gearbox is determined by opening the planetary gear\u2019s undercut grooves and developing a design method for the planetary gear hole parameters, as shown in Fig. 2. The lubrication method for the fan-driven gearbox is illustrated in Fig. 3. One of the oil separators is equipped with eight nozzles. The torsion support, input shaft, and output shaft that exert minimal influences on heat production and heat transfer are omitted from the model due to the complex internal structure of the fan-driven gearbox. Considering that the fan-driven gearbox adopts a five-way shunt structure and is structurally symmetrical, a one-fifth gearbox is adopted for modeling in the CFD simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001226_acpee48638.2020.9136520-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001226_acpee48638.2020.9136520-Figure5-1.png",
        "caption": "Fig. 5. The diagram of effect analysis for conductivity defects length on the electric field distribution in both composite insulator strings",
        "texts": [
            " The calculation results showed that the electric field at the high voltage end of insulator and the defect increase when the conductive defects existed in the insulator, and the longer the defect length was, the more the electric field increased. When the defect length was 0.6m, the electric field at the high voltage end of insulator string reached to 4226.5 kV/m, it is about 7 times of the value of the insulator without defects. The electric field at the defect was 610.5kV/m. b) Defects in the Both Insulator Strings: When the conductive defects exist at the both of the high voltage ends of two insulator strings, the diagram of influence analysis for conductive defects length on the electric field distribution is shown in Fig.5. In Fig. 5, the parameters L1 and L2 are the lengths of the defects. 1580 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on July 27,2020 at 03:12:24 UTC from IEEE Xplore. Restrictions apply. The calculation results of distribution of composite insulator strings with conductive defect are shown in Fig. 6, and the length of the defect section was set to L1=L2=0.6m. The calculation results of insulator electric field with conductive defects of different lengths at high voltage end are shown in Table III and Table IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001368_speedam48782.2020.9161859-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001368_speedam48782.2020.9161859-Figure2-1.png",
        "caption": "Fig. 2. Side view of induction motor",
        "texts": [
            " Downloaded on August 29,2020 at 16:24:07 UTC from IEEE Xplore. Restrictions apply. III. EXPERIMENTAL APPARATUS A. Induction motor To test the motor, one microphone was placed in front of the motor, in line with its shaft, and three accelerometers are placed in significant positions: 1) over the supporting base, 2) in line with the shaft and 3) at the frame back. The top view of the acoustic test rig is reported in Fig. 1. The side view of the test rig, showing the same sensors, is reported in Fig. 2: The characteristic parameters of the cage induction motor used for experiments are given in Tab.III-A: Specification Value Unit P 0,75 kW cos\u03d5 0,8 V\u0394 230 V Vy 380 V I\u0394 3,25 A Iy 1,88 A n 1395 rpm The electric drive is controlled with a Space-Vector Pulse-Width Modulation (SVPWM) technique, implemented in a field programmable gate array (FPGA) embedded in a control system NI-PXI built by National Instruments\u00ae. The control scheme is implemented in the Labview\u00aeFPGA environment. In particular, the FPGA generates PWM gate signals controlling the three-phase VSI feeding the motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000284_j.compeleceng.2019.106534-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000284_j.compeleceng.2019.106534-Figure5-1.png",
        "caption": "Fig. 5. Motor field-circuit coupling model.",
        "texts": [
            " Regarding the external circuit model, the induced electromotive force of stator winding is expressed with magnetic vector and the voltage balance equation of stator winding. It has the form as U s = R s i s + L \u03c3 d i s dt + l S (\u222b \u222b + \u2202 A z \u2202t d \u2212 \u222b \u222b \u2212 \u2202 A z \u2202t d ) (3) where U s and i s are the stator phase voltage and phase current, respectively, R s is the stator resistance, L \u03c3 is the stator winding leakage inductance, l is the axial length of iron core, S is the stator winding cross section area, and + and - are the total area of the positive and negative directions of the regional phase winding coils, respectively. Fig. 5 shows the motor field-circuit coupling model. As shown in the Fig. 5 , U A1 , U B1 , U C1 are torque winding phase voltages, L A1 \u03c3 , L B1 \u03c3 , L C1 \u03c3 are torque winding leakage inductances, W A1 + , W B1 + , W C1 + are torque winding beginnings, W A1- , W B1- , W C1- are torque winding endings; U A2 , U B2 , U C2 are suspension winding phase voltages, L A2 \u03c3 , L B2 \u03c3 , L C2 \u03c3 are suspension winding leakage inductances, W A2 + , W B2 + , W C2 + are suspension winding beginnings, W A2- , W B2- , W C2- are suspension winding endings. The torque generation principle of the BIM is similar to the ordinary induction motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000696_ilt-11-2019-0498-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000696_ilt-11-2019-0498-Figure1-1.png",
        "caption": "Figure 1 Schematic diagram of the plate-pin pair",
        "texts": [
            " (2019) established a thermal EHL model for crowned rollers and studied the effects of sliding-rolling ratio, load, crown radius and roller length on the lubrication state and temperature rise of in the contacts. However, to date, there has been no research work on the EHL of the plate-pin hinge pair in silent chains. Unlike rolling element bearings or roller chains, the thickness of plate is very thin. Thus, a model of finite line contact with narrow calculation domain should be used to solve the problem. In this study, this problem is solved by using a steadystate assumption. 2.1Mathematical models As shown in Figure 1, the contact between the plate and the pin is a conformal one: a is plate, b is pin and ra and rb are the radii of the plate and the pin shaft, respectively. The speed of the plate is ua, the speed of the pin is 0, l0 is the thickness of plate, ysk is the length of the rounded corner area at either side of the plate and rend is the radius of the rounded corner. The equivalent curvature radius r of the plate and the pin is calculated as follows: r \u00bc rarb= ra rb\u00f0 \u00de (1) In Equation (1), ra and rb are the radii of the plate interior hole and the pin, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000989_s40430-020-02392-5-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000989_s40430-020-02392-5-Figure5-1.png",
        "caption": "Fig. 5 Cylinders in contact",
        "texts": [
            " Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:298 Page 5 of 17 298 The first step toward calculating the solution of contact problems is determining the size and shape of the area of contact and distribution of normal pressure, giving stresses and leading toward the deformations. In this section, the common basic solutions in elastic contact problems are explained. In this work, the contact between two gear teeth is line contact; hence, this present estimation of induced forces is related to cylindrical pair. Figure\u00a05 shows cylinders in contact. When two cylindrical surfaces are pressed against each another with equal force F, a rectangular contact area of length L and width b is formed. Theoretically, this surface contact is a line type. The half width (b) of the rectangular contact area is presented by Eq.\u00a01. The pressure distribution in case of rectangular contact area is elliptical, and the maximum contact stress is given by Eq.\u00a02. In this work, the working stress is considered equal to maximum pressure ( Pmax) Since the gear and pinion are made up with the same materials, Poisson\u2019s ratio (v) and Young\u2019s modulus (E) are also same; hence, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001831_iccc49849.2020.9238795-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001831_iccc49849.2020.9238795-Figure1-1.png",
        "caption": "Fig. 1. UAVs equipped with directional antennas serve as BS in a finite area.",
        "texts": [
            " Downloaded on May 14,2021 at 17:22:14 UTC from IEEE Xplore. Restrictions apply. we utilize stochastic geometry to analyze the performance of finite UAV-enabled networks from a statistical view. We derive exact expressions of downlink coverage probability for directional and omnidirectional antennas, which provides general performance trends for setting altitudes and optimal sector angles of UAVs. We consider a wireless communication network where N UAVs equipped with BSs serving in a circular geographical area of radius Rc as shown in figure 1. The location of UAV i is denoted by [xi, yi, hi]. M GTs uniformly distribute on the ground of the finite area. The height of GTs is negligible compared to the altitude of UAVs. The UAVs reuse the same bandwidth B. Subchannels of each UAV are orthogonal. Subchannels are allocated to GTs based on the method described in [9]. The horizontal location of UAV i is denoted by [xi, yi]. The distance between a GT and the projection of UAV i on the ground is denoted by ri. The UAVs are equipped with directional antennas of which the half beamwidth is denoted as 2\u03b8, with \u03b8 \u2208 (0, \u03c02 )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001659_icuas48674.2020.9213887-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001659_icuas48674.2020.9213887-Figure2-1.png",
        "caption": "Fig. 2. The quadrotor reference frames.",
        "texts": [
            " The quadrotor propulsion and steering system consists of four independent motors and propellers with fixed pitch, where each pair of diagonal propellers rotate in opposite directions, to avoid yaw torque during roll and pitch movements. Hence, the system dynamics are mainly governed by the thrust and torque applied by the individual actuation (motor and propeller) units [23], [24]. Let us define a body-fixed frame Bi = {eBxi ,eByi ,eBzi } attached to the i\u2212 th vehicle\u2019s center of gravity OBi , as shown in Fig. 2, and an inertial frame I = {eIx ,eIy ,eIz} located at a fixed position OI within the workspace of the quadrotors. Following standard modeling techniques [24], the dynamic model of each vehicle may be derived from the general Newton-Euler motion equations of a 6-DOF rigid body subject to external forces and torques, as follows: I p\u0307Bi = IvBi mi I v\u0307Bi = IRBiFBi JBi \u03c9\u0307Bi = MBi (1) where IpBi = [ IxBi IyBi IzBi ]T , IvBi =[ IvxBi IvyBi IvzBi ]T are the position and linear velocity vectors of the i\u2212 th vehicle w",
            " for more details): FBi = FMi +Fdi +Fgi Mi = MMi +Mdi (2) where: \u2022 Fdi = \u2212Cd,F BiRI \u2225\u2225IvBi \u2212 IvwBi \u2225\u2225(IvBi \u2212 IvwBi ) are the drag forces with Cd,F denoting the drag force coefficient and IvwBi the wind velocity vector; \u2022 Mdi =\u2212Cd,M \u2223\u2223\u03c9\u0307Bi \u2223\u2223 \u03c9\u0307Bi are the drag moments with Cd,M denoting the drag moment coefficient; \u2022 Fgi = mi BiRI [ 0 0 ge ]T is the gravity vector with ge denoting the gravitational acceleration; \u2022 FMi = [ 0 0 \u2212(T1 +T2 +T3 +T4) ]T is the motor thrust vector; \u2022 MMi =  (T4\u2212T2) lm (T1\u2212T3) lm (\u2212M1 +M2\u2212M3 +M4) T is the motor torque vector and lm is the projected distance w.r.t the center of mass of each motor as shown in Fig.2; \u2022 Ti =CT \u03c9i 2, i = 1 . . .4 denote the thrust of each individ- ual thruster, CT > 0 is the thrust coefficient and \u03c9i is the rotational speed of i\u2212 th thruster; \u2022 Mi =CQ\u03c9i 2, i= 1 . . .4 denote the reaction torque of each individual thruster and CQ is the torque coefficient. In this work, we assume that the low-level control of each quadrotor is realized by a set of cascaded PID controllers [24]. More specifically, we distinguish: i) an inner loop which is responsible for the attitude, yaw rate and vertical velocity control of the vehicle and ii) an outer loop responsible for the horizontal velocity, heading and altitude control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001586_978-981-15-5463-6_82-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001586_978-981-15-5463-6_82-Figure1-1.png",
        "caption": "Fig. 1 Experimental setup",
        "texts": [],
        "surrounding_texts": [
            "This experimental work was carried out on machinery vibration simulator machine. This machine includes a single-stage bevel gearbox attached to a three-phase induction motor, with a load system. The driver gear has 27 teeth and driven gear has 18 teeth. The experimental setup also comprises a computer-aided DAQ. During the experimental investigation, the characteristics of the vibration signal were evaluated and analyzed in three different conditions of gearbox, i.e., healthy tooth gearbox, broken tooth gearbox, and missing tooth gearbox under the speed of 25 Hz with three specific loading conditions. Figures 1 and 2 show the experimental setup and faulty gearset. Total 3 samples were taken for experimental investigation where, 1 sample for the healthy condition, 1 sample for missing condition, and 1 sample for the chipped condition were collected for 30 s under the three different load conditions (0, 2, and 4 lb) of the gearbox with the speed of 15, 25, and 35 Hz. Figures 3 and 4 show the proposed fault diagnosis method and the raw vibration response for three different cases of the gearbox."
        ]
    },
    {
        "image_filename": "designv11_63_0002120_s00170-020-06571-5-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002120_s00170-020-06571-5-Figure13-1.png",
        "caption": "Fig. 13 Rolling test process",
        "texts": [
            " Table 3 Contact attribute targets Parameter type Drive side Coast side Coordinate of mean contact point in tooth width direction h /mm 0.000 0.000 Coordinate of mean contact point in tooth height direction v /mm 10.650 12.650 Orientation angle of contact path \u03b8/\u00b0 23.000 Long axis length of contact ellipse at mean contact point L/mm 10.000 Transmission error at the mesh-in and mesh-out point TEmax/\u03bcrad 90.000 The rolling test between the grinding wheel and pinion was conducted on the rolling inspection machine, shown in Fig. 13. Figure 14 shows the real contact patterns during the rolling test, which are consistent with the calculated contact patterns from the TCA software in Fig. 11, proving the feasibility and correctness of the proposed method. In this study, the drive sides are the wheel convex flank and the pinion concave flank, and the coast sides are the wheel concave flank and the pinion convex flank. If the drive sides and coast sides interchange in some situations, this proposed method is applicable as well. The contact attributes targets in this study are not optimal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002721_tec.2021.3074818-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002721_tec.2021.3074818-Figure2-1.png",
        "caption": "Fig. 2 Superposition based on a single stator tooth of a FRPM machine with stator-PM. (a) Stator with Ns teeth. (b) Stator with 1st tooth only. (c) Stator with ith tooth only.",
        "texts": [
            " Moreover, other characteristics can also be modeled by the superposition method based on either a single stator tooth or a single rotor tooth, including back-EMF, cogging torque, UMP, etc. as expressed: \ud835\udf11(\ud835\udf03, \ud835\udefc) = \ud835\udc53(\ud835\udf03) \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (2) where, \ud835\udf11(\ud835\udf03, \ud835\udefc) is the characteristic related to B(, ), \ud835\udc53(\ud835\udf03) is the function related to PM-excited MMF, and \ud835\udc54(\ud835\udf03, \ud835\udefc) is the function related to air-gap permeance. Setting the FRPM machine shown in Fig. 1(b) as an example, the mechanism of superposition of a single stator tooth can be described as follows. The open-circuit characteristics of the machine with \ud835\udc41\ud835\udc60 stator teeth [Fig. 2(a)] can be modeled by the superposition of those having a stator with single tooth only. It should be noted that the positions of the single stator tooth and rotor teeth, i.e., Figs. 2(b)-2(c) remain the same as Fig. 2(a). So does the rotor topology. Then, the characteristic of the machine with 1st stator tooth only [Fig. 2(b)] is: \ud835\udf111(\ud835\udf03, \ud835\udefc) = \ud835\udc531(\ud835\udf03) \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (3) Then, given the space angle between the first tooth and the \ud835\udc56th tooth being 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1), the characteristic of the machine having the rotor with \ud835\udc56th tooth only, i.e., Fig. 2(c), is written as \ud835\udf11\ud835\udc56(\ud835\udf03, \ud835\udefc) = \ud835\udc53\ud835\udc56 [\ud835\udf03 + 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1)] \u2219 \ud835\udc54(\ud835\udf03, \ud835\udefc) (4) Therefore, the characteristic of the machine having \ud835\udc41\ud835\udc60 stator teeth, i.e. Fig. 2(a), can be expressed by the superposition of those having a stator with single tooth only: Similarly, the mechanism of superposition based on a single rotor tooth can be described as follows. The open-circuit characteristic of the machine with \ud835\udc41\ud835\udc5f rotor teeth, shown in Fig. 3(a), can be modeled by the superposition of those having a rotor with single tooth only. The position of each single rotor tooth, i.e., Figs. 3(b)-3(c), remains the same as Figs. 3(a). Besides, the stator topology also remains the same",
            " 4(a) compares the air-gap flux density by both superposition method and original 10-polerotor machine, where \u2018S (analytical)\u2019 means the analytical predicted result by superposition, \u2018S (FEA)\u2019 means the FEApredicted result by superposition, and \u2018Original\u2019 means the FEA-predicted result of original 12/10 FRPM machine. The key parameters and dimensions of the proposed machines are listed in Table I, including stator outer/inner diameters (Dso/Dsi), air-gap length (\u03b4), stack length (La), PM remanence (Br) and PM coercivity (Hc). Clearly, a good agreement can be obtained. In addition, the air-gap flux density also can be modeled by the superposition of single stator tooth. As we can see from Fig. 2, the space angle between every two teeth is 2\ud835\udf0b \ud835\udc41\ud835\udc60 . Then, according to (4) and (9), the flux density of the FRPM machine with \ud835\udc56th tooth only is: \ud835\udc35\ud835\udc60\ud835\udc50\ud835\udc56(\ud835\udf03, \ud835\udefc) = \u2211 \u2211 \ud835\udc35\ud835\udc5b3\ud835\udc5a4 sin{(\ud835\udc5b3 \u00b1\ud835\udc5a4\ud835\udc41\ud835\udc5f)\ud835\udf03 + 2\ud835\udc5b3\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 \u221e \ud835\udc5a4=0 \u221e \ud835\udc5b3=1 1) \u00b1 \ud835\udc41\ud835\udc5f\ud835\udc5a4\ud835\udefc} (17) By using the same simplification as (13)-(14), the flux density equation obtained by this superposition method also has the same formation as (9). To verify above analyses, Fig. 5(a) compares the air-gap flux density calculated by both superposition methods and original 12-slot-stator machine",
            " Cogging torque in FRPM machines can be expressed as [26]: \ud835\udc47\ud835\udc50\ud835\udc5c\ud835\udc54(\ud835\udefc) = \u2212 \ud835\udf15 \ud835\udf15\ud835\udefc [ 1 2\ud835\udf070 \u222b\ud835\udc39\ud835\udc5d\ud835\udc5a 2(\ud835\udf03) \ud835\udeec2(\ud835\udf03, \ud835\udefc)\ud835\udc51\ud835\udc49] (19) where, \ud835\udf070 and \ud835\udc49 are permeability and volume of the air gap, respectively. As can be seen, cogging torque expression relates to both PM-MMF and air-gap permenace, which means it can be modeled by the superposition method as well. The general expression of cogging torque [26], [27] is: \ud835\udc47\ud835\udc50\ud835\udc5c\ud835\udc54(\ud835\udefc) = \u2211 \ud835\udc47\ud835\udc581 \u221e \ud835\udc581=1 sin[\ud835\udc581\ud835\udc3f\ud835\udc36\ud835\udc40(\ud835\udc41\ud835\udc60, \ud835\udc41\ud835\udc5f)\ud835\udefc] (20) where, \ud835\udc47\ud835\udc581 is the amplitude of cogging torque, \ud835\udc3f\ud835\udc36\ud835\udc40 is the least common multiple, \ud835\udc581 is harmonic order. As we can see from Fig. 2, the space angle between every two teeth is 2\ud835\udf0b \ud835\udc41\ud835\udc60 . Then, according to (4) and (20), the cogging torque component of the FRPM machine with \ud835\udc56th tooth only is: \ud835\udc47\ud835\udc50\ud835\udc60\ud835\udc56(\ud835\udefc) = \u2211 \ud835\udc47\ud835\udc582 \u221e \ud835\udc582=1 sin {\ud835\udc582\ud835\udc41\ud835\udc5f[\ud835\udefc + 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1)]} (21) Then, by using the same simplification as equations (12)-(13), the cogging torque equation obtained by this superposition method can be simplified as: \ud835\udc47\ud835\udc50\ud835\udc5c\ud835\udc54(\ud835\udefc)= \u2211 \ud835\udc47\ud835\udc582 sin\ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b sin \ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b \ud835\udc41\ud835\udc60 \u221e \ud835\udc582=1 sin (\ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udefc- \ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b \ud835\udc41\ud835\udc60 +\ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b) (22) As can be seen, (22) is meaningful only when the denominator sin \ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b \ud835\udc41\ud835\udc60 \u2261 0 since sin\ud835\udc582\ud835\udc41\ud835\udc5f\ud835\udf0b is always 0",
            " According to [22], the \ud835\udc39\ud835\udc5d\ud835\udc5a 2(\ud835\udf03) and \ud835\udeec2(\ud835\udf03, \ud835\udefc) are \ud835\udc39\ud835\udc5d\ud835\udc5a 2 (\ud835\udf03) = F\ud835\udc5d\ud835\udc5a0+ \u2211 F\ud835\udc5d\ud835\udc5a\ud835\udc5b4cos(\ud835\udc5b4Ns\u03b8) (29) \u221e \ud835\udc5b4=1 \ud835\udeec2(\ud835\udf03, \ud835\udefc) = \ud835\udeec0 + \u2211 \ud835\udeec\ud835\udc5a5 cos\ud835\udc5a5\ud835\udc41\ud835\udc5f(\ud835\udf03 + \ud835\udefc) \u221e \ud835\udc5a5=1 (30) where, F\ud835\udc5d\ud835\udc5a0, F\ud835\udc5d\ud835\udc5a\ud835\udc5b4, \ud835\udeec0 and \ud835\udeec\ud835\udc5a5 are Fourier coefficients. Substitute equations (29) and (30) into (27), and thus the UMP due to PM only can be calculated (set the \ud835\udc65 component as an example). \ud835\udc43\ud835\udc65(\ud835\udefc) = { \u2211 \u2211 \ud835\udc43\ud835\udc5b4\ud835\udc5a5 cos ( \u221e \ud835\udc5b4=1 \u221e \ud835\udc5a5=1 \ud835\udc5a5\ud835\udc41\ud835\udc5f\ud835\udefc),\ud835\udc5a5\ud835\udc41\ud835\udc5f = \ud835\udc5b4Ns \u00b1 1 0, elsewhere (31) where, \ud835\udc43\ud835\udc5b4\ud835\udc5a5 is Fourier coefficient. As can be seen, the UMP expression relates to both PMMMF and air-gap permenace, which means it can be modeled by the superposition method. For example, as can be seen from Fig. 2, the space angle between every two teeth is 2\ud835\udf0b \ud835\udc41\ud835\udc60 . Then, according to (3), (4) and (27), the \ud835\udc65 component in FRPM machine with \ud835\udc56th tooth only is: \ud835\udc43\ud835\udc65\ud835\udc56(\ud835\udefc) = { \u2211 \ud835\udc43\ud835\udc5b5cos\ud835\udc5b5 2\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1), \ud835\udc5b5 = 1 \u221e \ud835\udc5b5=1 \u2211 \u2211 \ud835\udc43\ud835\udc5b5\ud835\udc5a6 cos[\ud835\udc5a6\ud835\udc41\ud835\udc5f\ud835\udefc \u2212 \u221e \ud835\udc5b5=1 \u221e \ud835\udc5a6=1 2\ud835\udc5b5\ud835\udf0b \ud835\udc41\ud835\udc60 (\ud835\udc56 \u2212 1)], \ud835\udc5a6\ud835\udc41\ud835\udc5f = \ud835\udc5b5 \u00b1 1 0, elsewhere (32) Therefore, according to (5), the total UMP (\ud835\udc65 component) can be obtained by the superposition of those having a stator with single tooth only. Besides, with the similar simplification of equations (12) to (13), the UMP can be expressed as: \ud835\udc43\ud835\udc65(\ud835\udefc) = { 0, \ud835\udc5b5 = 1 \u2211 \u2211 \ud835\udc43\ud835\udc5b5\ud835\udc5a6 \u2219 sin\ud835\udc5b5\ud835\udf0b sin \ud835\udc5b5\ud835\udf0b \ud835\udc41\ud835\udc60 cos [\ud835\udc5a6\ud835\udc41\ud835\udc5f\ud835\udefc \u2212 \ud835\udc5b5\ud835\udf0b(\ud835\udc41\ud835\udc60 \u2212 1) \ud835\udc41\ud835\udc60 ] \u221e \ud835\udc5b5=1 \u221e \ud835\udc5a6=1 , \ud835\udc5a6\ud835\udc41\ud835\udc5f = \ud835\udc5b5 \u00b1 1 0, elsewhere (33) As can be seen from (33), the UMP exists only when sin \ud835\udc5b5\ud835\udf0b \ud835\udc41\ud835\udc60 \u2261 0 since sin\ud835\udc5b5\ud835\udf0b is always 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002010_j.tws.2020.107334-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002010_j.tws.2020.107334-Figure5-1.png",
        "caption": "Fig. 5. Static simulation results of the cylindrical shell.",
        "texts": [
            " Firstly, the static analysis of the cylindrical shell with actuation force is performed with finite element method. The driving force (stress) of the SMA tube is assumed as 40 MPa. The boundary condition of the cylindrical shell is simply supported at both ends. The contact form of the actuator and the inner wall of the cylindrical shell is set as \u2018no separation\u2019; the contact form of the actuator and the circumferential sliding block is set as \u2018bonded\u2019. The static deformation simulation results can be obtained as shown in Fig. 5. In order to analyze the stress situation at the contact position between the actuator and the cylindrical shell more intuitively, we set the 12 o\u2019clock position in the circumferential direction of the cylindrical shell as 0\u25e6. The stress on the inner wall of the cylindrical shell at 30\u25e6, 45\u25e6 and 60\u25e6 is extracted respectively, as shown in Fig. 6. One can find from Fig. 6 that the peak value of the stress occurs at the contact position between the actuator and the inner wall of the cylindrical shell"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002900_09544054211023625-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002900_09544054211023625-Figure4-1.png",
        "caption": "Figure 4. Basic setup of the assembly error evaluation items for spiral bevel and hypoid gears.",
        "texts": [
            " The hypoid generator setting modification considering the noise factors such as assembly error, SGE, thermal error, wear, and clamping error in actual machining environment has been rarely reported in literature.19,21 In consideration here of Z, assembly error and SGEs are considered as the main noise factors to develop functional relationships with respect to hypoid generator settings. As mentioned in Artoni et al.,23 their function with hypoid generator settings can be represented as Z1 = P,G,E,a\u00bd 7!X Z1\u00f0 \u00de \u00f06\u00de The detailed computation process can refer to Artoni et al.23 As indicated in equation (6) and Figure 4, the assembly error generally includes four evaluation items, namely Z1= [P, G, E, a]. Its detail definition is developed in Ding et al.,26 Shao et al.,48 and Simon.51 The relations of SGE with respect to hypoid generator settings in modification have been developed in Ding et al.20 There are totally 33 evaluation items and their function is represented as Z2 = SGE1, ,SGE33\u00bd 7!M Z2\u00f0 \u00de \u00f07\u00de The detailed solution about SGEs can refer to Artoni et al.23 Their definitions and descriptions are omitted for brief here"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure17-1.png",
        "caption": "Fig. 17 Schematic diagram of the evaluation points",
        "texts": [
            " Based on the actual structure of the double-layer box, fixed support constraints were set at the bolt holes of the outer box, and dynamic load was applied to the input bearing seat, star wheel carrier, and output bearing seat. The boundary conditions of the box are shown in Fig.\u00a016. The box body was subjected to load and vibration response was generated at different positions. Four points were selected at the output bearing seat, input bearing seat, and outer box foot as the evaluation points of the vibration response, which were named as evaluation point I-1\u20134, II-1\u20134, and III-1\u20134, respectively (Fig.\u00a017). Figures\u00a018, 19, and 20 show the time and frequency domain response diagrams for the different evaluation points. The peak acceleration and frequency of each vibrating body corresponding to Figs.\u00a018, 19, and 20 are shown in Tables\u00a07, 8, and 9, respectively. According to Figs.\u00a018, 19, and 20: 1 3 1. In the transient response results obtained through FEM, the peak acceleration of the output bearing seat was between 6.59 and 8.34\u00a0m s\u22122, that of the input bearing seat was between 64.91 and 69.14\u00a0m s\u22122, and that of the outer box was between 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002411_iros45743.2020.9340935-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002411_iros45743.2020.9340935-Figure12-1.png",
        "caption": "Fig. 12: (a) Confined space test. Inchworm robots can pass through the 8cmhigh confined space while the quadruped robot is too large. (b) Obstacle climbing test. The quadruped robot could climb the 5mm-high obstacle while the inchworm robot could not.",
        "texts": [
            " Then, following the gait control strategy we designed, the quadruped robot stood up and began walking. The locomotion ability of the quadruped robot was tested on both the nylon cloth surface and an ordinary wooden floor. The frequency of the trot gait walking is fixed as 2 Hz and the average velocities are 0.12m/s and 0.1m/s, respectively. Compared with the single inchworm robot under the same conditions, the maximum velocity of the quadruped configuration is 70% faster. We performed the confined space passing experiment on both the inchworm robot and quadruped robot (Fig. 12(a)). The result shows that the inchworm can pass through an 8cmhigh path while the quadruped robot cannot, demonstrating that the small size and flexibility of the inchworm robot is 11685 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 17,2021 at 20:54:10 UTC from IEEE Xplore. Restrictions apply. beneficial to the mobility of the bionic robotic system because the whole system can pass through confined spaces as a swarm of inchworm robots. The test of the climbing ability of the quadruped robot is shown in Fig. 12(b). The whole system can assemble into the quadruped robot configuration to climb a 5mm-high step while the individual inchworm robots could not. Even though a step height of 5mm is less than ideal for a quadruped robot of this size, it demonstrates the adaptability of the BioARS to rough terrain benefiting from the quadruped configuration. We designed a bionic assembly robotic system (BioARS) with inchworm robots as modules. The assembly configuration of the system is a quadruped robot. Onboard power and control systems are integrated to make the inchworm robot module untethered and flexible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001295_14484846.2020.1794522-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001295_14484846.2020.1794522-Figure2-1.png",
        "caption": "Figure 2. Laser drilling system schematic diagram.",
        "texts": [
            " The proposed methodology has been presented in Figure 1. Trepanning experiments are carried out on a 1.4 mm thick sheet of material Inconel-718 using solid-state 250 W, pulsed Nd: YAG laser machining system. In this system, the beam of laser light is delivered at work material sheet in normal to surface direction with the help of CNC controlled motion of delivery nozzle. The specification of the drilling system has been shown in Table 1. The schematic diagram of the laser trepan drilling system has been shown in Figure 2. Wide engineering applications of nickel-based superalloy Inconel-718 have been reported in the advanced technological fields like gas turbine components, marine equipment, medical equipment, and similar critical application. Compositions of Inconel-718 has been mentioned in Table 2. Moreover, the selected input parameters have been done based on past work done by various researchers. Several factors have confined the selection of input parameters like; specification of machine, the feasibility of variation in input parameter, and availability of resources"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003191_itec51675.2021.9490116-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003191_itec51675.2021.9490116-Figure3-1.png",
        "caption": "Fig. 3. Different reluctance rotor topologies [55]. (a) axially laminated rotor, (b) transversally laminated rotor, (c) simple salient rotor.",
        "texts": [],
        "surrounding_texts": [
            "The machine topology and drive system configuration of the proposed five-phase synchronous reluctance motor drive is introduced in this section. Fig. 1 shows cross-section, power converter, and drives topologies for the proposed fivephase synchronous reluctance motor. The five-phase symmetric distributed windings and simple salient reluctance rotor are employed in the proposed synchronous reluctance motor. The rotor structure selection will be discussed in the next section, which will show that the simple salient reluctance rotor is more suitable for the proposed five-phase synchronous reluctance motor because of its good coupling capability between fundamental and third harmonic MMF [51]-[52]. In addition, a full-bridge five-phase power electronic converter was employed to achieve vector control and harmonic current injection for proposed five-phase synchronous reluctance motor drives. The third harmonic current component can be utilized for average torque production. The electromagnetic torque of the proposed five-phase synchronous reluctance motor can be expressed [53] (1) where P is pole numbers, Ld1, Lq1 are fundamental d-axis and q-axis inductance, Ld3, Lq3 are third harmonics d-axis and qaxis inductance, id1, iq1 are d-axis and q-axis fundamental current, id3, iq3 are d-axis and q-axis third harmonic current. The first term of this equation represents the electromagnetic torque produced by fundamental current, which is similar to the torque production in a conventional three-phase synchronous reluctance motor. In addition, the third term represents electromagnetic torque produced by the third harmonic current only, which has a very small contribution and can be neglected. It should be noted that the second term indicates the interaction of fundamental and third harmonics current and MMF. This interaction can produce useful average torque because of the special coupling effect of the salient reluctance rotor. The torque capability can be improved and enhanced by utilizing the third harmonic injection. III. DESIGN AND VERIFICATION A. Rotor topology selection Rotor topology or structure plays a key role in order to improve additional torque capability in the proposed fivephase SynRM. Generally, the topologies of reluctance rotor can be classfied into three types, including salient pole, multilayer flux barriers, and axially laminated rotor. Both multilayer flux barriers and axially laminated rotors can achieve a higher saliency ratio than simple salient pole rotors. The multi-layer flux barrier reluctance rotor is an attractive option in conventional three-phase synchronous reluctance motor since it compromises saliency ratio and manufacturing complexity. However, in order to utilize the effect of the third harmonic current and MMF, the simple salient rotor is more suitable in the proposed five-phase SynRM. The influence of different rotor structures on the torque capability of fivephase SynRM has been investigated in [49]. In addition, the simple salient reluctance rotor is also attractive for high-speed traction applications. Therefore, the simple salient reluctance rotor is employed in the proposed design and analysis to meet electromagnetic and mechanical design requirements. B. Performance verification In this section, the finite element analysis was employed to predicted electromagnetic performance and characteristics of 1 1 1 1 13 1 3 1 3 3 3 3 3 ( ) 2 ( )5 3( )2 2 q d d q q d d q e q d d q L L i i L i i i iP T L L i i \u2212 \u2212 \u2212   =   + \u2212   Authorized licensed use limited to: Hoseo Univ. Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply. the proposed five-phase synchronous reluctance motor with harmonic current injection. The key parameters are summarized in TABLE 2. Based on the analysis in the previous section, a simple salient pole reluctance rotor is more suitable for the proposed five-phase SynRM with harmonic current injection. Fig. 4 shows flux distribution of proposed five-phase SynRM with simple salient pole reluctance rotor exited by d-axis and q-axis current respectively. Fig. 5 shows flux density under peak power and continuous power operation, respectively. Authorized licensed use limited to: Hoseo Univ. Downloaded on September 02,2021 at 04:21:35 UTC from IEEE Xplore. Restrictions apply. Fig. 6 shows the exciting current waveform of symmetric five-phase windings. Fig. 7 shows the phase current waveform with a different ratio of third harmonic current. It should be pointed out that the current (rms value) of third harmonic and fundamental components keep unchanged to maintain the same copper losses. Fig. 8 shows torque variation with different ratios of third harmonic and fundamental current. There is an optimal ratio of third harmonic and fundamental current. The torque achieves its maximum value when the ratio is around 0.4 or 0.5. Fig. 9 shows variation of average torque with different current angles under peak and continuous power operation. The torque capability improvement with and without third harmonic current injection was summarized in TABLE 3. It can be seen that the torque capability can be enhanced by 9.6% (under peak power operation) and 12.4% (under continuous power operation) respectively. The torque improvement of peak power operation is lower than continuous power operation because the torque improvement is sensitive to the saturation level of the stator and rotor core. The effect of harmonic current can be further improved by using advanced lamination materials. IV. CONCLUSION A robust five-phase synchronous reluctance motor have been presented in this paper. The machine topology, torque production mechanism, drive system configuration, rotor structure selection were discussed. The finite element method was used to predict the electromagnetic performance. The average torque can be improved by around 9% ~12% utilizing a third harmonic current injection. The experimental verifications will be included in future publications. The proposed design has a robust rotor structure, which is suitable for high-speed traction applications. Additionally, the proposed five-phase synchronous reluctance motor avoids some concerns and challenges in conventional IPM machines such as uncontrolled generation and demagnetization risk since the permanent magnets are eliminated. Consequently, the proposed five-phase synchronous reluctance motor offers an appealing permanent magnet-free option for future electric vehicle drivetrain."
        ]
    },
    {
        "image_filename": "designv11_63_0000015_aim.2019.8868337-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000015_aim.2019.8868337-Figure2-1.png",
        "caption": "Figure 2. Mechanical analysis of overcoming obstacle",
        "texts": [
            " Moreover, the effect of wheelbase on the performance of crossing obstacle ability is analyzed, which provides a theoretical basis for the wheelbase changing strategy during crossing obstacle. 978-1-7281-2493-3/19/$31.00 \u00a92019 IEEE 1329 The concept of variable-wheelbase chassis is illustrated in Figure 1. The longitudinal position of the middle axle can actively change relative to the frame of chassis. Each wheel can be driven independently and the independent suspensions of each wheel have the same stiffness. The structural parameters of the chassis are explained in Table 1. Figure 2 shows the force acting on each part of the wheel in the process of crossing the obstacle. It also illustrates the mechanical analysis of overcoming obstacle for front axle (Figure 2a and Figure 2b), middle axle (Figure 2c) and rear axle (Figure 2d). ( 1,2,3)iN i  is the support force of the ground acting on axles,  is the ground adhesion coefficient. The equilibrium equation of force in each direction and the moment balance equation for the front axle can be obtained from figure 2(a): 1 2 3(sin cos )N N N W      (1) 1 2 3( sin -cos ) 0N N N       (2) 1 2 1 3( ) ( ) 0rN N r l N r l Wa        (3) The relationship between the support force acting on each wheel and the spring deflection can be described by the following equation: 1 1 2 2 3 3 (sin cos )N k N k N k             (4) In order to obtain the deflection relationship between the suspension springs from the geometric parameters of the vehicle by proportional method, the following assumptions are made on the mathematical model:  The mounting position of the suspension spring on the vehicle body is in a line",
            " This evaluating method is based on the assumption that the initial state is the hardest situation in the process of overcoming obstacle. Therefore, it is necessary to verify the correction of above assumption. Assume that the front wheels have lifted a certain height during the process of crossing the obstacle. Equations of the force and moment balance are established in this condition, and then find the hardest situation for overcoming obstacles by increasing s from zero to h. The force and moment balance equations of the front axle can be obtained by analyzing Fig 2. (b): 1 2 3(sin cos )N N N W      (8) 1 2 3( sin -cos ) 0N N M       (9)     1 2 1 3 0 ( ) cos ( )sec ( ) cos ( tg )cos 0 rN N r s l l l N r s l W a h                    (10) The relationship between the support force acting on the front wheels and the deflection of the suspension springs can be expressed as:  1 1 2 2 3 3 sin( ) cos( ) (cos sin ) (cos sin ) N k N k N k                          (11) The deflection relationship between the suspension springs is:    1 1 2 1 3 1l l l l l l l tg        (12) Using equations (8)-(12), After some calculation, the expressions obtained are:   2 1 1 1 1 1 1 2 21 1 (1 / ) (1 )cos [ 1 sec 1 1 cos 1 sec (1 )sec (1 ) (1 ) 1 cos cos ]o l l l tg la kl G l l W G l l l lr s AB l l l l l hl lr A A tg l l l l                                                                (13) where cos sin sin cos sin cos cos sin sin( ) cos( ) A B C D C B                               (14) sin sin sin s h l r r s r h l l l r                 (15) Taking the same method, analysis equations of the middle axle can be obtained from Figure 2 (c): 1 2 3(sin cos )N N N W      (16) 1 2 3( sin cos ) 0N N N        (17)       1 2 1 2 3 0 sin cos ( )sec sin cos cos sin 0 N r l N l l rN rN W l a h                        (18) The relationship between the support force on the wheel and the deflection of the corresponding suspension spring can be expressed as:   1 1 2 2 3 3 (cos sin ) sin( ) cos( ) (cos sin ) N k N k N k                          (19) Deflection relationship between springs:    1 1 2 1 3 1l l l l l l l tg        (20) Combining equations (16), (17), (18), (19) and (20):               2 2 1 1 2 1 1 1 2 1 cos cos sin 2 (1 )cos 2 sec 0 o kl B l l r ABl l l tg G W l a h l l l l l l C r                                  (21) cos sin sin cos sin cos cos sin sin( ) cos( ) A B C G C B                               (22)  sin 1 sin r l    (23) Similarly, analysis equations of the rear axle can be obtained from Figure 2 (d): 1 2 3(sin cos )N N N W      (24) 1 2 3( sin cos ) 0N N N        (25)       1 2 1 3 0 rsin cos rsin cos sec r cos h sin 0 N l N l l N W l a                      (26) The relationship between the support force on the wheel and the deflection of the corresponding suspension spring can be expressed as:   1 1 2 2 3 3 (cos sin ) (cos sin ) sin( ) cos( ) N k N k N k                          (27) Deflection relationship between springs:  1 1 2 1 3 1+l l l l l l tg        (28) Solving the equations (24)-(28):               2 1 2 2 1 1 2 1 1 1 1 1 / cos sin / 2 / 1 cos 2 sin cos sec 1 cos sec 2 =0 oA a l h l l l kl AB l l r l l tg W l AB l l Gl A l l r                                 (29) To verify the assumption that the initial time of crossing obstacle is the most difficult situation in the process, the parameters in table 1 are substituted into equation (13)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001609_0021998320960531-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001609_0021998320960531-Figure6-1.png",
        "caption": "Figure 6. (a) Numerical model for beam 3point bending test, a fine mesh was applied at the vicinity of indenter beam interface. (b) Longitudinal dimensions of beam and indenter.",
        "texts": [
            " The flexural properties of parent SrPP material were investigated following ASTM standard D 790 \u2013 03 using the Instron machine with loading nose and supports of radii 5.0 0.1mm so that no local stress concentration occurred. The 3-point bending tests on SrPPSB were performed using screw driven uniaxial Instron 4045 testing machine with crosshead speed of 0.5mm/min at room temperature according to standard ASTM D 7264 as followed by other authors.27 The radii of loading nose and supports were 6.0 0.1mm so that no local stress concentration occurred. A schematic with dimensions and actual test setup is shown in Figure 6(b). As the consolidated SrPP material consists of 0/90 plain woven tapes so in plane material properties i.e. 1\u20132 direction are considered same. Tensile tests were performed at cross head speed of 1mm/min and because of woven symmetry in direction 1 and 2 elastic modulus was measured to be E1\u00bcE2\u00bc 3.03GPa, similar to the measured by Gabor et al.28 E\u00bc 2.95GPa. Quasi static compressive modulus Ec was found to be 3.75GPa and a peak compressive stress r1,2\u00bc 28.56MPa, while tensile strength was measured rY1,2 \u00bc 141",
            " When corrugated SrPPSB are subjected to the out of plane loading the deformation phenomenon and associated analytical models are very complex due to a large number of material and structural parameters. So commercial FEA simulation packages along with the proposed constitutive material models are efficient tools to investigate the mechanical behaviour of complex structures. This can save the extensive experimental work, cost and time to optimize the design parameters. The numerical model as shown in Figure 6 was developed in commercial FE package ABAQUS 6.14 explicit for its excellent nonlinear ability. The indenter with 12mm diameter was modelled as discrete rigid component and the beam was modelled with linear reduced integration solid elements C3DR8. Further details about the elastic-plastic material model used can be found in.29 The numerical model is validated with experimental results in our previous paper as shown in Figure 7.30 We fabricated self-reinforced polypropylene (SrPP) sandwich beams (SrPPSB) with non-symmetric mass distribution between core and face sheet (FS) and three point bending tests were performed on SrPPSB to investigate the effects of mass distribution on their flexural properties and energy absorption capacity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001146_s00773-020-00743-4-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001146_s00773-020-00743-4-Figure2-1.png",
        "caption": "Fig. 2 Coordinate system",
        "texts": [
            " Also, a model of the subject ship had undergone capsize due to broachingto in the high-speed region during the model experiments carried out by Umeda and Hamamoto [16] at the National Research Institute of Fisheries Engineering (NRIFE), Japan. It is desirable to use a simpler but still reasonable model for a nonlinear system because the increase in the number of degrees of freedom makes the nonlinear dynamical system more difficult to be analyzed. Hence, the authors use a 4 DoF surge-sway-yaw-roll model [17, 18] whose coordinate systems are as shown in Fig.\u00a02, which consists of two coordinate systems: a) A wave fixed system with its origin located on the calm water level, above a wave trough with axis in the direction of wave travel, \u03b6 axis pointing downwards, and \u03b7 axis along the wave trough. b) An upright body-fixed system with its origin at the center of gravity of the ship, and the x, y, and z axes pointing towards the bow direction, towards the starboard, and downwards, respectively. The PD controlled system is defined using the state vector, x, and the parameter vector, p, as follows: 1 3 And the state equations are: where The symbols used in Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001206_s12206-020-0604-7-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001206_s12206-020-0604-7-Figure1-1.png",
        "caption": "Fig. 1. Reduced-order finite element model of blisk.",
        "texts": [
            " The frequency is usually limited to a specific excitation order at a narrow range, thus the calculation only involves a small amount of dominant modal. According to the actual situation, the accidents or failures of aeroengine often occur in blades and less occur in disk, therefore the blade is regarded as mistuned substructure and disk is regarded as tuned substructure. The internal DOFs of disk are reduced, the contact DOFs between blades and disk are regarded as constraint modal, and internal DOFs of blades are regarded as the main modal. Then, the reduced-order finite element model of mistuned blisk is established by RSCMSM in Fig. 1. The main parameters are shown in Table 1. To verify the computational efficiency of RSCMSM, the number of nodes and elements extracted, which compared with those of CSCMSM and HFFEM, it shows in Table 2. As seen in Table 2, the number of nodes and elements of CSCMSM relative to HFFEM is reduced 62.79 % and 71.27 %, the number of nodes and elements of RSCMSM relative to HFFEM is reduced 82.35 % and 83.67 %, it means the number of nodes and elements of RSCMSM relative to CSCMSM is reduced 52.56 % and 43"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002419_iros45743.2020.9341316-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002419_iros45743.2020.9341316-Figure1-1.png",
        "caption": "Fig. 1. (a)-(b) Examples of repositioning and reorienting an object grasped by a gripper. (c)-(d) Examples of reorienting and manipulating heavy objects.",
        "texts": [
            " Nevertheless, it is still difficult to utilize these hands for complex manipulation tasks except for pick-and-place grasping tasks. To increase ability of the robotic hands and manipulators to grasp and manipulate a wider range of objects, environment can be effectively exploited. Successful accomplishment of many tasks involve exploitation of contacts of the grasped objects with the environment [1]. In general, environment contact can be exploited in two main ways: (1) environment contacts as external forces for re-positioning and reorienting an object which is already grasped by a robotic hand or gripper (Fig. 1-a,b), [2],[3],[4] (2) environment contacts as essential supports for reorienting a heavy object or moving it from one position to another (Fig. 1-c,d). The former task can enable simple affordable grippers to do dexterous manipulation [2]. The latter task can enable the robotic arms to manipulate heavy or large objects effectively. For moving heavy or large objects, a common strategy used (by humans) is to lift the object by pivoting on one edge and then, move it, while maintaining contact with the ground. This strategy exploits the contact of the object with the ground so that the arms do not need to carry the full weight, yet the object can be manipulated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002567_j.anucene.2021.108241-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002567_j.anucene.2021.108241-Figure2-1.png",
        "caption": "Fig. 2. Distribution of lamination strips along angular periphery in actual ALIP.",
        "texts": [
            " With advent of high power computation facilities, analyses of ALIP using finite element method (FEM) based techniques have also been reported. This paper presents the FEM based analysis of \u2018\u2018end-effects\u201d due to discontinuous magnetic circuit in amedium-gap ALIP by comparing its performance against a continuous magnetic circuit ALIP. This medium-gap ALIP is designed to pump liquid sodium at temperatures up to 450 C. ALIP has circular winding in the form of disks which are housed in lamination stacks. These lamination stacks are distributed around the angular perimeter of ALIP as shown in Fig. 2. The flux produced due to current flowing in winding travels through these stacks and enters the annular duct radially and penetrates liquid sodium and enters the central magnetic core made up of laminated magnetic steel. When flux enters the duct region, its distribution is primarily symmetrical around the axis of the duct. Thus, model of ALIP can be made in a two-dimensional axisymmetric space. Leakage of flux from one laminated stator stack to another and consequent decrease in flux penetrating the duct is not included in such an axisymmetric model",
            " Additionally, there is no leakage of magnetic flux at entry and exit ends. Due to noninclusion of ends extending out of inductor core, a CMC-ALIP becomes shorter in length than a DMC-ALIP. Comparison of results predicted by two models will help in evaluating the influence of \u2018\u2018end-effects\u201d in a real-world ALIP which always has end-effects. In numerical models of both CMC-ALIP and DMC-ALIP the distribution of lamination strips around the periphery of stator is assumed to be uniform (Fig. 6) in contrast to an actual ALIP where the stacks are discretely distributed (Fig. 2). Computations were carried out at fixed terminal voltage of 360 V, 50 Hz and sodium temperature of 450 C. Sodium flow rate was also varied and various characteristics of ALIP were obtained. In order to have a better understanding of phenomena taking place, variations in magnetic flux density, current density and volume force density have also been compared at rated flow and with static sodium. The variation in magnitude of radial component of magnetic flux density (Br) for CMC and DMC ALIPs is depicted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001787_ecce44975.2020.9235832-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001787_ecce44975.2020.9235832-Figure5-1.png",
        "caption": "Fig. 5. Flux density field line plots with (a) zero and (b) rated wound-field excitation.",
        "texts": [
            "945 Harmonic leakage flux coefficient \u03c4d 1.175 Number of magnet poles 32 Number of wound-field poles 16 Number of rotor field slots and field coils 16 Stator inner diameter 530 mm PM rotor outer diameter 655 mm Wound rotor outer diameter 740 mm Axial stack length 114 mm Airgap length 1.8 mm RMS stator current density 6.0 A/mm2 Rated (maximum) field current density 6.0 A/mm2 core losses are evaluated in Section VII. The flux density field line plots of this machine without and with wound-field excitation as illustrated in Fig. 5 clearly show the action of hybrid rotor field excitation. FE analysis is used to determine the d-axis flux linkages to calculate the induced voltage versus field current of the machine. This result is shown in Fig. 6, where a \u00b120% voltage variation is obtained for the rated 400 V machine in Fig. 3. 1442 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 24,2021 at 15:17:54 UTC from IEEE Xplore. Restrictions apply. In this section we briefly consider the FE modelling of the inductances of the generator, and a solution method is provided for performance estimation at specific power and reactive power operating points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001164_s12541-020-00369-x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001164_s12541-020-00369-x-Figure1-1.png",
        "caption": "Fig. 1 Three-stage lever scheme (a) and picture (b)",
        "texts": [
            " Therefore, our work focuses on developing a system model that can predict and optimize the PEA\u2019s performance. We considered the piezoelectric material parameters and their dependence on the external loading conditions. The main goal of this paper is to describe and evaluate a model of PEMs which are mounted in a practical mechanism. The amplification mechanism consists of three amplification stages, mounted on one another. Each stage acts like a lever with nonlinear properties, because of the non-constant position of the contact points between the levers, as shown in Fig.\u00a01. These levers convert linear movements of 50\u00a0\u03bcm to angular movement of 40\u00b0. The linear displacement of three parallel piezostacks produces the rotational movement of lever #1, and lever #2 and finally causes the rotation of the output axis. The model in this paper describes and analyzes the amplification mechanism and its influence on the angular movement of the last amplification stage, as well as the load\u2019s influence on the PEM and the overall characterization of these electro-mechanical systems",
            " Measurements were performed on a laboratory model. The experimental results are reported in this section and compared to the model predictions. Three identical discrete actuator stacks were fitted in tandem in the mechanism set-up. The piezoelectric stacks were mechanically connected in series, but the electrical driving voltage was connected in parallel. Since the piezoelectric stacks were connected mechanically in series, their displacements were added together to yield a significant enhancement of the total displacement, as seen in Fig.\u00a01. In order to increase system accuracy and to reduce displacement losses and parasitic motion, a mechanism with adequate stiffness to handle the piezostack was designed and fabricated, as described in [48]. The dynamic response of the model was observed by measuring the displacement and the feeding voltage characteristics, while keeping the load conditions unchanged. The PEA was loaded by a weight mounted at the end of the third amplification stage. The simulation model has some unknown parameters, such as viscous friction and pre adjustment; these parameters were hand-tuned in order to optimize the model and make it suitable for different modes of excitation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001035_tmech.2020.2997132-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001035_tmech.2020.2997132-Figure1-1.png",
        "caption": "Fig. 1. An unmanned aerial vehicle (a) uses a three-axis miniaturized gimbal system (b) to stabilize its onboard camera. (Picture courtesy of [1] and [2]).",
        "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. I. INTRODUCTION Inertially stabilized multi-axis gimbal systems have been widely used on a variety of mobile platforms, such as unmanned aerial vehicles (UAVs), for payload (e.g., cameras) stabilization. As the host platform provides the necessary mobility in 3D space, it is required to stabilize the attitude of the payload, motivating the use of a 3-axis gimbal system as shown in Fig. 1. A gimbal system usually consists of three joints that are driven by direct-drive motors. The payload attitude relative to the inertial reference frame is measured by two MEMS inertial measurement units (IMUs), one on the payload (i.e., the endpoint IMU) and the other on the host vehicle (i.e., the vehicle IMU). In addition to these, other sensors such as magnetometer and GPS may also be available on the host vehicle. In this paper, we consider a type of compact and cost-effective gimbal systems, in which the endpoint IMU contains a rate gyro and an accelermeter only",
            " For better visualization, both plots are shown in log scale. It can be seen that the actual axis-angle error remains well below our predicted error bound, which verifies our proposed error model. Additionally, both the actual axis-angle error and its error bound grow linearly (with slope equal to 1 in the log scale) with respect to the joint control error, which agrees with the prediction in (27). In this section, the proposed calibration algorithm is implemented and verified on a DJI Mavic Pro onboard gimbal system (see Fig. 1). The X , Y and Z axes of the endpoint gyro are respectively aligned with the gimbal roll, pitch, and yaw directions. In the experiment, each to-be-calibrated joint axis moves along the profile provided in Fig. 2; the calibration data are collected every 1 degree. Instead of global basis functions, the encoder nonlinearity map is fitted using firstorder interpolation due to its reduced requirements of memory and computation. Using the calibrated axis orientation and encoder nonlinearity map, the reference attitude is computed by the gimbal forward kinematics in (1) and fused with the gyro measurements with a multiplicative Kalman filter [37]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001869_s11223-020-00217-3-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001869_s11223-020-00217-3-Figure8-1.png",
        "caption": "Fig. 8. Locations of element Y , X , D, and E within the wheel model.",
        "texts": [
            "0035 s is significantly smoother in the peak and valley vicinities than with 0.0140 or 0.007 s, and can therefore ensure high accuracy of peak and valley values\u2019 calculation. Thus, 0.0035 s was selected as the sampling time in the modal transient dynamic analysis of the wheel. 1601 transient response points were calculated, covering a time period of 5.6 s and about 40 cycles. The modes with modal frequencies below 1000 Hz were used for the calculation. A damping value of 0.02 was set for each mode. Figure 8 shows four representative elements, Y , X , D, and E, with element Y being sensitive to F y , element X to Fx , and elements D and E to both (Fig. 6). Figure 9 shows their transient normal Y and normal X strain histories. It can be seen that the transient response of each element stabilized after about 2 s. The normal Y strain histories of elements Y and E in the left part of the finite element model of the wheel had much larger steady amplitudes than the corresponding normal X strain histories",
            " The corresponding formula for any strain history is ij ij k FEA k k k n t P P t( ) ( ) , , , 1 (7) where n is the total number of load cases, ij t( ) is the strain or stress history (such as normal Y strain history) to be determined, P tk ( ) is the kth load-time history, P tFEA k, ( ) is the static force corresponding to P tk ( ), and ij k, is the strain or stress corresponding to ij t( ), which is obtained via the static finite element analysis applying only PFEA k, . According to formula (7), the normal Y strain history of element Y and the normal X strain history of element X (Fig. 8) can be calculated in the fatigue analysis using the following formulas: nY nY y FEA y y nY x FEA x xt P F t P F t( ) ( ) ( ), , , , , (8) nX nX y FEA y y nX x FEA x xt P F t P F t( ) ( ) ( ), , , , , (9) where F tx ( ) and F ty ( ) are the same as Fx in formula (1) and F y in formula (2), respectively, nY t( ) is the normal Y strain history of element Y , nX t( ) is the normal X strain history of element X , PFEA y, is the static force corresponding to F ty ( ), let P FFEA y, 0 2344.4 N with F0 being the amplitude of F ty ( ) or F tx ( ), nY y, is the normal Y strain of element Y obtained via static finite element analysis applying only PFEA x, , nX y, is the normal X strain of element X obtained via static finite element analysis applying only PFEA y, , PFEA x, is the static force corresponding to F tx ( ), P FFEA x, 0 2344"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure1-1.png",
        "caption": "Fig. 1. RMxprt model of the PMSM.",
        "texts": [
            " For the design of PMSM, the toolbox RMxprt (Rotational Machine Expert) can be used. This tool allows its user to design pre-defined motor patterns. As mentioned above, the simulation time is one of the core problems of the co-simulation. Via RMxprt and motor design (winding scheme/slot to pole ratio etc.), the PMSM can be shared into its symmetry parts. This means, instead of doing the simulation of the complete motor, only one of its symmetric components is simulated and then the complete motor behaviour can be extrapolated. Fig. 1 shows the RMxprt model of the studied machine. When the RMxprt model was analysed, it was exported to the two-dimensional and three-dimensional model using the ANSYS Maxwell Software. In this case, the machine has been divided into four equal parts and only one part is represented on 2D and 3D model as shown in Fig. 2 and Fig. 3. The geometrical of 2D and 3D model are depicted in Fig. 2 and Fig. 3. 682 Authorized licensed use limited to: Raytheon Technologies. Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001180_1077546320940872-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001180_1077546320940872-Figure2-1.png",
        "caption": "Figure 2. Overhead crane model.",
        "texts": [
            " The theoretical findings were experimentally validated using a prototype crane. This study\u2019s major finding demonstrates that the new technique succeeded in reducing the maneuvering time by 12.9%\u2013 21.25% with ZV at the motion end. Moreover, the results show the insensitivity of the proposed shaper to variations in system parameters using the ZVD shaper. An overhead crane is modeled as a single pendulum of constant cable length attached to a movable trolley (jib) restrained to move in a rectilinear motion along the x-axis, as shown in Figure 2. The cable is carrying a lumped mass representing the payload that swings in the xy-plane. The jib is moving according to a predefined input acceleration function. For simplicity, system nonlinearities such as motor dead zone, are ignored in the crane modeling. Generally, input dead zones of practical servomotors are essential (Yang et al., 2020). Neglecting these nonlinearities can reduce the control performance and in some cases, lead to instability. Lagrange\u2019s equation is implemented to obtain the equation of motion of the system, which is given by L\u20ac\u03b8 \u00fe g sin \u03b8 \u00bc \u20acu cos \u03b8; \u03b8\u00f00\u00de \u00bc _\u03b8\u00f00\u00de \u00bc 0 (1) where L is the cable length, g is the gravity constant, and \u20acu is the jib acceleration function"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure13-1.png",
        "caption": "Figure 13. First and second eigenmode (78.17 Hz and 80.67 Hz).",
        "texts": [],
        "surrounding_texts": [
            "Composite materials, finite element analysis, automotive engineering, vibrational behavior, design optimization Date received: 21 March 2021; accepted: 5 May 2021"
        ]
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure5-1.png",
        "caption": "Fig. 5. A detailed view of the mesh quality in the 3-D model of Permanent Magnet Synchronous Machine.",
        "texts": [
            " The geometrical of 2D and 3D model are depicted in Fig. 2 and Fig. 3. 682 Authorized licensed use limited to: Raytheon Technologies. Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. During the FE simulation of permanent magnet synchronous machine, the meshing is essential process which is done to descretize the geometry developed into small number of parts called cells. A quarter geometry of the PMSM is subdivided into triangular elements (1000 elements). As obvious from Fig. 4 and Fig. 5 both models give relatively the detailed view of the mesh quality in the 2D and 3D model of the PMSM. Fig. 6 shows the electromagnetic torque calculated with the 2D and 3D FE models: The torque obtained from this model is also shown in Fig. 6. It can be seen that the shape of the electromagnetic torque differs noticeably from the others. The blue curve (3D) obviously lies much closer to the measurement. It is because the machine is relatively short and the edge effects are (not considered in 2D) not negligible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001334_ccssp49278.2020.9151834-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001334_ccssp49278.2020.9151834-Figure1-1.png",
        "caption": "Fig. 1: Rolling bearing",
        "texts": [
            "00 \u00a92020 IEEE Authorized licensed use limited to: Middlesex University. Downloaded on August 03,2020 at 03:19:30 UTC from IEEE Xplore. Restrictions apply. IV. Furthermore, the description of the proposed method is reported in section V. Then, the results on both simulated and measured vibration signals are showed in sections VI and VII. Finally, section VIII concludes this work. The bearing consists fundamentally of both outer and inner raceways, balls and the cage which keeps unchanged the distance between the balls (see Fig. 1). The bearing faults in the literature can be classified according to two main categories: single-point defects and generalized roughness. The single-point fault which appears either as a very small hole or a missing material on the outer and the inner raceways, produces mechanical shockwaves every time the rolling elements pass over it. Consequently, this will excite the frequencies of natural mechanical resonance in the motor. Its occurrence rate is equal to one of the characteristic fault frequencies (inner or outer race ways and balls faults)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000795_ec-03-2019-0083-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000795_ec-03-2019-0083-Figure8-1.png",
        "caption": "Figure 8. Diagrammatic drawing of meshing gear pair at end plane",
        "texts": [
            ", 2013): e t\u00f0 \u00de \u00bc em \u00fe ea cos v t \u00fe w\u00f0 \u00de (15) Where em shows the average term, ea represents the fluctuation and w refers to the initial phase angle. In general, gear backlash is caused by gear lubrication and errors in manufacturing, as well as installation, which would make the meshing contact of gear pair change between contact and separation state. Hence, backlash should be considered in the system. Thus, the relative deformation f(x) of the meshing teeth can be described as: f x\u00f0 \u00de \u00bc x b x> b 0 jxj# b x\u00fe b x< b 8>< >: (16) Where b refers to the half of gear backlash and x is the relative displacement of the teeth along the mesh line. Figure 8 shows the diagrammatic drawing of the meshing gear pair at the end surface. In Figure 8, rai and ri (i = 1, 2) represent the radius of addendum and pitch circles, respectively. B1 and B2 are the departure and terminal points of the theoretical meshing line, respectively. A1 and A2 are the departure and terminal points of the actual meshing line, respectively. Meantime, l is the length of the actual meshing line. l2 refers to the distance from pitch point P to ending pointA2 of the actual meshing line. Item Value from Ishikawa/energy method Maximum stiffness in single teeth meshing (109 N/m) 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002348_s12206-021-0133-z-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002348_s12206-021-0133-z-Figure9-1.png",
        "caption": "Fig. 9. Corners detection of the pattern fixed to a spherical joint.",
        "texts": [
            " A flat pattern is attached to one of the spherical joints, and then we proceed to make several photographs of it (30 in this case) in different positions, Fig. 8. The vertex space of the i -th spherical joint attached to the mobile platform, M iV , is a sphere with center at ( ), ,c i i i iS a b c . Superscript c indicates that these coordinates are relative to the camera reference frame. Additionally, the origin of the coordinate system of the calibration pattern, represented by point ( ), , , ,, ,c c c c i f i f i f i fO x y z , as seen from the camera in the photograph f , Fig. 9, belongs to the surface of the sphere, which is described by the well-known equation ( ) ( ) ( )2 2 2 2 , , , c c c i f i i f i i f i ix a y b z c r\u2212 + \u2212 + \u2212 = (5) where ir is the radius of the sphere, and parameters ia , ib and ic are the coordinates of the sphere\u2019s center. The process of fitting a sphere to a set of points consist on minimizing the differences between the radii, ir r\u2212 . Hence, the objective function is given by ( ) ( ) ( )2 2 2 , , ,( ) c c c i i f i f i f id x a y b z c r= \u2212 + \u2212 + \u2212 \u2212u "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure11-1.png",
        "caption": "Fig. 11. Flux density distribution (Maxwell 3D).",
        "texts": [
            " Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. Waveforms of stator winding currents and the induced voltage obtained from Ansys Maxwell software of the both models (2D and 3D) are compared in Fig. 8 and Fig. 9. Using Maxwell 2D and Maxwell 3D, the magnetic flux density distribution may be properly compared to evaluate the influence of the both types of the model using ANSYS Maxwell software [3]. Waveforms of the magnetic field calculated in Maxwell 2D and 3D are shown in Fig. 10 and Fig. 11 for time t = 0.2 s and at the speed of 1500 rpm. In Fig. 10 the maximum flux density value is 2.7069 T. It can be seen from Fig. 11 that the flux density value is 2.6465 T. Small differences between the 2D and 3D solutions in the ANSYS Maxwell software calculated of the permanent magnet synchronous motor are primarily caused by influence of the motor end windings and the slot skewing. From the comparison of currents waveforms, electromagnetic torque and flux density distribution it is clear that the differences between results of two dimensional and three dimensional models are relatively small but must be considered. In case of preliminary, calculation the two dimensional model gives satisfying results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000120_012090-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000120_012090-Figure2-1.png",
        "caption": "Figure 2. Location (a) and typical design (b) of the orbital joints of the MC-21 aircraft compartment",
        "texts": [
            " One of the most common types of joints is riveting, bolt riveting and bolt joints, which make up to 90% of the total number of joints of parts [1, 2, 3], as well as in all aircraft. When assembling the MS-21 aircraft, impact riveting technology is prohibited, and in the absence of approaches for the assembly of parts, bolted and riveted bolt connections are used. In particular, most commonly used when connecting the panels along the longitudinal joints in the process of assembling the compartments (Fig. 1) and when connecting the panels along the orbital joints in the process of assembling the fuselage (Fig. 2), bolt-ribbing and, less often, bolted joints. ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10.1088/1757-899X/632/1/012090 The design of the stud bolt and bolt connections varies depending on the type of rod and bolt head, the size of the head, the material used for the fixing element, and the method of installation in the hole. Examples of rivet and bolt connections are shown in Figure 3 [7]. ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012090 IOP Publishing doi:10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002118_s00521-020-05554-7-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002118_s00521-020-05554-7-Figure2-1.png",
        "caption": "Fig. 2 Kinematic diagram of the twin-double pendulum (robot and symbolic human leg)",
        "texts": [
            " There is no actuator in the wrist joint, which is only used to inhibit the capacity of human movement when transferring the robot\u2019s entire load to the ground. The human leg and robot body were modeled as twin-double pendulums. In the twin-double pendulum mechanism, the dimensions are the same for each pendulum, which are connected by a spring located between them. To determine the error between the robot and the human leg, the amount of shortening or elongation in the springs was used. The twin-double pendulum mechanism\u2019s kinematic diagram is shown in Fig. 2. One of the twin-double pendulums in the diagram is the human leg and the other is the robot system following it. The models of this twin-double pendulum system were coupled with the spring components representing the sensors during the simulation studies, and their interaction was ensured. The human leg is shown in blue and the robot parts are colored in red. The angles of the limbs are indicated by hi, their mass mi, their length li, and their center of gravity Gi. Through constructing the dynamic model, the torque expressions needed in the twin-double pendulum system are obtained for the joints of the human leg mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000100_j.jaerosci.2019.105482-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000100_j.jaerosci.2019.105482-Figure1-1.png",
        "caption": "Fig. 1. Scheme of the planar oblique impact of a homogeneous sphere dropped vertically without initial spin on an infinitely massive inclined plane. a) Center of mass velocities and impact and rebound angles; b) restitution and friction impulses.",
        "texts": [
            ", 2018), where the oblique impact of fly ash particles with stainless surfaces is described, is of particular interest because the values of the normal coefficient of restitution are below 0.50 and the values of the tangential coefficient of restitution, using the aforementioned semi-analytical AFR closure (Gorham & Kharaz, 2000; Kharaz et al., 2001; Wu et al., 2009), vary significantly with the impact angle. The current study indicates that, in all cases, the reported experimental data can be reasonably described by the IFR analytical model using a unique set of the values of the coefficients of restitution and friction. Fig. 1 shows schematics for the experimental set up used in the study of fly ash particle oblique impact with stainless surfaces (Xie et al., 2018). Ideally, this corresponds to the planar oblique impact of homogeneous rigid spheres of mass m, radius R, and inertia moment I (\u00bc (2/5)mR2) on a rough and infinitely massive stationary plane. Let u and U the pre-rebound velocities of the mass center of the sphere and its contact point with the plane, respectively, and v and V, the respective post-rebound velocities",
            " In sliding regime of impact, this can be expressed in terms of the coefficient of sliding friction, \u03bc: For collisions in stick regime, there is no disposal of a constitutive equation equivalent to the Amontons-Coulomb law for sliding friction (Herbst et al., 2000; Luding, 2008; M\u00fcller & P\u20acoschel, 2012; Schwager et al., 2008). Based on percussion dynamics, a reasonable proposal is (Dom\ufffdenech-Carb\ufffdo, 2016): Jf \ufffd t \u00bc 2 5 m\u00f0U \ufffd t\u00de sign\u00bdU \ufffd t\ufffd (9) For the case of the oblique rebound of a homogeneous sphere on an infinitely massive plane without initial spin, the normal impulse will be, introducing the angle of incidence, \u03b3, as described in Fig. 1: Jn\u00bc \u00f01\u00fe en\u00demu cos \u03b3 (10) In turn, the net tangential impulse acting in sliding regime will be: Jt \u00bc 2 7 \u00f01\u00fe et\u00demu sin \u03b3 \u03bc\u00f01\u00fe en\u00demu cos \u03b3 (11) Whereas the net tangential impulse acting on the sphere in sticking regime will be: Jt \u00bc \ufffd 2 5 \u00fe 2 7 \u00f01\u00fe et\u00de \ufffd mu sin \u03b3 (12) In cases where relatively large indentation can occur, there is possibility of the incorporation of the rolling friction effects into impact dynamics (Pishkenari, Rad, & Shad, 2017). In line with Iwashita and Oda (1998) and with Orlando and Shen (2010), there are two possible regimes of impact in this regard: rolling and nonrolling",
            " (Dom\ufffdenech-Carb\ufffdo, 2019): Krf \u00bc 2 5 Rmu sin \u03b3 (14) Combining equations (10)\u2013(14) it is possible to describe four theoretical regimes of impact: sliding plus rolling, sliding plus nonrolling, sticking plus rolling and sticking plus nonrolling (Iwashita & Oda, 1998; Orlando & Shen, 2010; Dom\ufffdenech-Carb\ufffdo, 2019). According to the precedent IFR formulation, the post impact normal and tangential components of the center of mass velocity, v, will be independent of the rolling/nonrolling regime. The normal component will be vn \u00bc enun, i.e., introducing the rebound angle \u03b4 A. Dom\ufffdenech-Carb\ufffdo and C. Dom\ufffdenech-Casas\u00fas Journal of Aerosol Science 140 (2020) 105482 (see Fig. 1): vn\u00bc v cos \u03b4 \u00bc enu cos \u03b3 (15) The corresponding expressions for the tangential component in sliding and sticking regimes derived from Eqs. (10)\u2013(12) are respectively. vt \u00bc v sin \u03b4 \u00bc u sin \u03b3 2 7 \u00f01\u00fe et\u00deu sin \u03b3 \u03bc\u00f01\u00fe en\u00deu cos \u03b3 (16) vt \u00bc v sin \u03b4\u00bc u sin \u03b3 \ufffd 2 5 \u00fe 2 7 \ufffd \u00f01\u00fe et\u00deu sin \u03b3 (17) In turn, the post rebound angular velocity and the tangential component of the velocity of the contact point will be influenced by the rolling friction effects. By introducing the non-dimensional coefficient of rolling friction, \u03b7 \u00bc \u03c1/R, Eqs (10)\u2013(14) yield: Sliding plus rolling regime: Vt \u00bc etu sin \u03b3 \ufffd 7 2 \u03bc 5 2 \u03b7 \ufffd \u00f01\u00fe en\u00deu cos \u03b3 (18) Sticking plus rolling regime: Vt \u00bc \ufffd 7 5 \u00fe et \ufffd u sin \u03b3 \u00fe 5 2 \u03b7\u00f01\u00fe en\u00deu cos \u03b3 (19) Sticking plus nonrolling regime: Vt \u00bc \ufffd 2 5 \u00fe et \ufffd u sin \u03b3 (20) Sliding plus nonrolling regime: Vt\u00bc\u00f01 et\u00deu sin \u03b3 7 2 \u03bc\u00f01\u00fe en\u00deu cos \u03b3 (21) The postrebound angle defined by the center of mass velocities, \u03b4, will be independent on the rolling/nonrolling effects"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002037_j.engfailanal.2020.105126-Figure22-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002037_j.engfailanal.2020.105126-Figure22-1.png",
        "caption": "Fig. 22. Stubs and legs.",
        "texts": [
            " 2) If subjected to traction, the load percentage recommended, as follows: Pulling force in the angle / (stress strain S.F.) Where the tensile capacity is the minimum between: - NCAP = tensile capacity based on the net section [37]. - RCAP = tensile capacity based on connection breakdown. - SCAP = Cutting capacity of the connection. - BCAP = Crush and tear capacity. Structural Model (TOWER Program) The finite element model, is for both suspension and retention in the Fig. 20, Fig. 21; besides, the structural model of the TO type towers generated in the Fig. 22, Fig. 23, Fig. 24, Fig. 25, Fig. 26, Fig. 27 and Fig. 28. In the Fig. 14, the suspension tower is modeled with finite element; each part of the tower has been built for the validation of the loads according to the section 3. In the Fig. 15, the material for retention tower has built, with the material influence for the calculations and constraints as elastic module, yield stress, steel type and components. Each component of the tower has been described in the Fig. 16 with the tower top. Fig. 17 for both upper arms, it applies to arm 1 and 2, Fig. 18 with the lowest arm of the tower, Fig. 19 in the upper body between arm 1 and 2, Fig. 20 for the body structure between arm 2 and 3, Fig. 21 and Fig. 22. Fig. 14. Tower model in finite element. R.M. Arias Vela\u0301squez and J.V. Mej\u00eda Lara Engineering Failure Analysis 121 (2021) 105126 The stage division of the ice disaster is composed of four stages: Preconditions (associated to the pre-failure period), steady-state progression and multiple failures (during disaster period) and restoration (post-failure period) [38]; recent papers have indicated effective resilience enhancement framework for towers installed, however, the recommendations for new assets are not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002176_0142331220981430-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002176_0142331220981430-Figure1-1.png",
        "caption": "Figure 1. Geometry for TA engagement.",
        "texts": [
            " Self-loops are not permitted in the communication topology, which means (i, i) is not allowed. A directed route from node i to node j is defined as a sequence of paths, (i, k1), (k1, k2),..,(kl, j), with different nodes km,m= 1, 2, ::, l. A graph contains a spanning tree means that there exists a directed rooted tree that utilizes a directed path starting from the root vertex to every other vertex. In this part, we consider the target-attacker scenario where N attackers pursue one maneuvering target in Figure 1, where Ai and T denote the i-th attacker and the target. Ri is the relative distance between the target and the i-th attacker, and Vi is a positive constant that describes the initial speed of the ith attacker. The terms gi and gT are the heading angles of the attacker and the target. li is the i-th attacker\u2019s LOS angle in the inertial reference frame. ji is the bearing angle between the direction of the i-th attacker\u2019s velocity and the i-th LOS, and the term fi is the bearing angle between the direction of the target\u2019s velocity and the i-th LOS"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000350_s1061830919100036-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000350_s1061830919100036-Figure4-1.png",
        "caption": "Fig. 4. Dependence of normal component of resulting field at surface on coordinate in plane for various values",
        "texts": [
            " For the farthest possible point from the ends (near ), the normal component is near zero, a fact that is consistent with the obvious result of solving the problem about an infinite magnetic cylinder placed in a constant external field directed along the cylinder axis. Figure 3 shows the graphs of for various cylinder lengths. With the length increasing, one can observe a trend of the value of the function decreasing near Based on the above, one can conclude that when placed in such an external field, a cylinder whose longitudinal size is at least twice its diameter can be considered as infinite when measurements are taken near its midpoint. Figure 4 illustrates the results of calculating the normal component of the resulting field in the cross section provided that the external field is of the form According to the exact solu- tion of the corresponding problem about an infinite cylinder, [1]. It can be seen from the figure that the points fit on this curve, albeit a slight difference noticeable near the extrema, which can be explained by the influence of the cylinder butt ends. Figure 5 shows the graphs of the same dependences 0,H 0,H 0H=H H C ( ){ } ;, 1; 1 i jN Ni j mk m k= = A 1,\u00a02,\u00a03i = 1,\u00a02,\u00a03j = jk jkS\u2208r jkS 1,\u00a02,\u00a03j = 1,\u00a0 ,\u00a0 jk N= \u2026 ( ){ } 11,1 1 N mm m= A nH \u03a9 ( )nH z\u03a9 ( )0 00,\u00a00,\u00a0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002639_wiecon-ece52138.2020.9397983-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002639_wiecon-ece52138.2020.9397983-Figure2-1.png",
        "caption": "Fig. 2. The single-track model's free body diagram.",
        "texts": [
            " Improving vehicle efficiency and control by adding steering system compliance parameters and accuracy has increased to 12 percent and 8% for steady state and transient response, respectively [7], Model predictive control (MPC) is useful for developing an electrical vehicle path-tracking controller [8], as well as vehicle yaw rate control using PID control technology [10]. II. TRACK MODEL The single-track model is a classic model that represents the most basic model for accurately predicting the lateral dynamics of a road vehicle. It is useful in non-extreme conditions, such as when the longitudinal speed is constant and the sideslip angle is minimal. On the same axle, the two tires are clumped together, resulting in one front and one rear tire, as seen in figure 2. Table 1 contains descriptions of the symbols on the figure. The equations of motion based on the free-body diagram in Figure 2 can be written using Newton's second law, and the model can be implemented using Matlab software and the motion equations. 20 20 IE EE In te rn at io na l W om en in E ng in ee rin g (W IE ) C on fe re nc e on E le ct ric al a nd C om pu te r E ng in ee rin g (W IE C O N -E C E) | 97 8- 1- 66 54 -1 91 7- 8/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D Authorized licensed use limited to: California State University Fresno. Downloaded on June 24,2021 at 00:22:55 UTC from IEEE Xplore. Restrictions apply. To complete this project, a single-track Simulink model using Matlab should be implemented, as shown in Figure 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001719_biorob49111.2020.9224276-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001719_biorob49111.2020.9224276-Figure1-1.png",
        "caption": "Fig. 1: Robotic palpation system.",
        "texts": [
            " For instance, ultrasound imaging has been used to visualize the tongue while a subject is speaking, providing feedback for speech therapy [7]. Our work presented in this paper takes place in a research project that focuses on the development of a robotic assistance system for ultrasound elastography (USE). USE has been explored in medicine for the diagnosis of breast tumors [8], liver fibrosis at different stages [9] and prostate cancer [10]. In a previous work [11], we developed a robotic palpation system that consists of an ultrasound probe held by a 6 degrees of freedom (DOF) robot (see Fig. 1a). A force control law was proposed to automatically and continuously apply a quasi-static compression over the tissue with the ultrasound probe in order to obtain the pre- and post-compressed states of the tissue. A method was also proposed to estimate in real-time the elastic parameters as strain values of the observed tissue from the radio frequency (RF) data acquired by the ultrasound probe between two consecutive ultrasound images. In this previous work, the tissue elastic parameters was provided to the examiner by a visual feedback thanks to the display of a color-coded image that represents a 2D elasticity map of the tissue called elastogram",
            " The first mode allows the user to teleoperate the ultrasound probe for changing the ultrasound view while applying the autonomous palpation motion and the second mode provides to the user a force feedback reflecting the tissue elasticity. Experimental results of these assistance modes are then presented in section IV. We briefly recall our previous work presented in [11] that concerned the design of a robotic system for soft tissue palpation for real-time elastography imaging. A force control law was proposed to automatically and continuously apply a quasi-static compression over the tissue with an ultrasound probe attached to a 6-DOF robot (see Fig. 1a). This compression process is performed to obtain the preand post-compressed states of the tissue that are required to estimate its elastogram (elasticity map) from radio frequency (RF) data acquired by the ultrasound probe. To obtain this continuous and periodical deformation of the tissue, we designed a force control law that applies a desired force variation along the y-axis of the cartesian frame Fcp attached at the bottom of the ultrasound probe (see Fig. 1a) that we defined as, s\u2217f (k) = \u2206F 2 [ sin ( (4k \u2212 T )\u03c0 2T ) + 1 ] + F0, (1) where k is the discrete time and \u2206F is the amplitude of a sinusoidal function. T is the period of the desired force signal expressed in sample time and F0 is the initial desired force value. To minimize the force error ef = sf \u2212 s\u2217f (with sf being the measured force by a force sensor along the y-axis), an exponential decrease of ef was achieved by imposing the desired error variation of the error such as e\u0307\u2217f = \u2212\u03bbfef with \u03bbf being the force control gain",
            " (14) with \u03b5 being the scalar strain value estimated from the ROI, and it gives: k = AE L (17) where A is the region area of the virtual probe that senses the local strain of the tissue in the ROI and that corresponds in our case to the elliptical surface A = \u03c0\u03c3x\u03c3y of the Gaussian mask Gm. E is the Young\u2019s modulus that we set to the value E = 3kPa of healthy tissue and L is the original length of the virtual spring that corresponds to the width d of the rectangular contact surface between the real ultrasound probe and the tissue (see Fig. 1b). 2) Virtual probe control and force feedback: Fig. 3 illustrates the principle that consists in moving the ROI to follow the displacement of the user measured by the handler of the haptic device. If the user applies motion at the handler of the haptic device Fh, then the center of the ROI (uc, vc) is shifted with a displacement \u2206d \u2208 R 2 proportional to the displacement of the handler such that, \u2206d = S\u2206h, (18) where \u2206h \u2208 R 2 is the in-plane relative motion applied to the handler that is directly measured from the x and y translation components of the 4th column of the homogeneous matrix h0 Mh introduced in eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001375_speedam48782.2020.9161929-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001375_speedam48782.2020.9161929-Figure2-1.png",
        "caption": "Fig. 2 \u2013 The filling of flux barrier in PMaSRM: the novel approach with bonded magnets",
        "texts": [
            " Some prototypes of assisted reluctance machine both with NdFeB and ferrite bonded magnets have been prepared with the adoption of different manufacturing techniques [31]. In the proposed work the two different concepts concerning the use of ferrites magnets in assisted reluctance machine have been investigated. The first approach consists in the adoption of regular shape magnets, which take up a partial part of the volume of the flux barriers, as shown in Fig. 1. In the other case, the flux barriers are completely filled with bonded magnets, as reported in Fig. 2. The study began with the investigation of a particular SRM with a rotor geometry designed for the propose of lowering torque ripple [36-38]. The flux barriers shapes have been optimized by means of a multi-objected differential evolution (DE) algorithm [39-42]. Some geometric dimensions and rated parameters concern SRM are shown in TABLE I. 978-1-7281-7019-0/20/$31.00 \u00a92020 IEEE 670 Authorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on August 12,2020 at 02:43:59 UTC from IEEE Xplore. Restrictions apply. The obtained rotor geometry illustrated in Fig. 1 and Fig. 2, has been adopted for evaluations performed by means of finite element analysis (FEA). During FEA estimations three different cases have been compared: SRM (without permanent magnets), PMaSRM with parallelepiped shape ferrite sintered magnets and PMaSRM with ferrite bonded magnets. After careful simulation analysis, the appropriate prototype will be prepared. All proposed magnets have been characterized in the laboratory. Moreover, the involved prototype has been also built and tested in the laboratory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001997_isec49744.2020.9280736-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001997_isec49744.2020.9280736-Figure2-1.png",
        "caption": "Figure 2: Planar Pendulum. Adapted from [7].",
        "texts": [
            " In Afghanistan a night mission to save a Soldier on the ground proved to be fatal as an oscillating rescue hoist [resulted in] the Soldier and rescue specialist losing their lives. This project seeks to reduce the risks inherent in an unstable hover.\u201d [6] This event was caused by a wild swing in displacement angle and could possibly have been avoided. It is the goal of this research to discover and implement this solution as soon as possible to bring home all our Soldiers. In order to solve this complex problem, it must first be simplified down to its simplest level. A two-dimensional planar pendulum with length and bob with mass is that simplest level, as shown in Fig. 2 [7]. A few assumptions must be made before moving forward in the analysis. First, we assume that there is no drag caused by air resistance, and second, we assume that the pivot point is frictionless. From the planar pendulum shown in Fig. 2, it can be seen that the equation for the speed of the pendulum, in Eqn. (1), where , (1) where the terms , and represent the speed, cable length, and angular speed, respectively. These are important to note as they are going to be substituted in the equation for kinetic energy, , as shown in Eqn. (2), such that, . (2) The following step is to write the Lagrangian equation. (Eqn. 4), where we subtract the potential energy (Eqn. 3), of the system, , from the kinetic energy of the system in order to obtain the Lagrangian (Eqn"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000112_1077546319874921-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000112_1077546319874921-Figure8-1.png",
        "caption": "Figure 8. The four-point powertrain mounting system.",
        "texts": [
            " F \u00bc M \u20acqp \u00fe Xn i 2 6666664 1 0 0 0 1 0 0 0 1 0 zpi ypi zpi 0 xpi ypi xpi 0 3 7777775 Fmi (3) 0 BBBBBB@ \u20acx1 \u20acy1 \u20acz1 \u20acx2 \u20acy2 \u20acz3 1 CCCCCCA \u00bc 2 6666664 1 0 0 0 zp1 yp1 0 1 0 zp1 0 xp1 0 0 1 yp1 xp1 0 1 0 0 0 zp2 yp2 0 1 0 zp2 0 xp2 0 0 1 yp3 xp3 0 3 7777775 (4) where Fmi is the dynamic reaction force of the mount i in the vehicle coordinate system (VCS), and the xpi, ypi, and zpi represent the location of mount i in the powertrain coordinate system. The results are shown in Figure 7. Figure 3. The 13-degree of freedom vehicle dynamic model with a semi-active hydraulic damping strut. Table 1. Stiffness and damping. Percentage of the semi-active hydraulic damping strut sharing force (%) Dynamic stiffness (N/mm) Damping coefficient (N s/mm) 80 1035.8 41.01 The semi-active HDS can be regarded as a mount with stiffness and damping, and the new four-point powertrain mounting system is shown in Figure 8. A 13-DOF vehicle dynamic model is built in order to analyze the influence of semi-active HDS on vehicle vibration response and calculate the structural parameters of the semi-active HDS, which contains tires, unsprung mass, suspension, vehicle body, and the powertrain mounting system. The powertrain mounting system has six DOFs, the vehicle body has three DOFs, and the four unsprung mass have four vertical DOFs. The 13-DOF vehicle dynamic model with a semi-active HDS is shown in Figure 3. The equation of the 13-DOF vehicle dynamic model is (Wang, 2017)2 64 Mp Mb Mu 3 75 8>< >: \u20acqp \u20acqb \u20acqu 9>= >;\u00fe 2 64 C6\u00d76 11 C6\u00d73 12 06\u00d74 C3\u00d76 21 C3\u00d73 22 C3\u00d74 23 04\u00d76 C4\u00d73 32 C4\u00d74 33 3 75 8>< >: _qp _qb _qu 9>= >; \u00fe 2 64 K6\u00d76 11 K6\u00d73 12 06\u00d74 K3\u00d76 21 K3\u00d73 22 K3\u00d74 23 04\u00d76 K4\u00d73 32 K4\u00d74 33 3 75 8>< >: qp qb qu 9>= >; \u00bc 8>< >: Fp 0 0 9>= >; (5) Here, the displacements of CoGs for the powertrain, vehicle body, and unsprung mass are defined as qp, qb, and qu, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000083_mees.2019.8896375-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000083_mees.2019.8896375-Figure1-1.png",
        "caption": "Fig. 1. External rotor generator.",
        "texts": [
            " PURPOSE OF RESEARCH For wind turbine generators, the rotational speed of which is n = 100 \u00f7 200 rpm, to carry out a comparative analysis of two magnetic systems of permanent magnet generators: a multi-pole external rotor generator (ERG) and a generator with a magnetic gearbox (MGM). For given dimensions of the electric generator (external diameter of the stator and axial length) of two devices, to determine the configuration and dimensions of the magnetic system of the gearbox, at which the maximum power value is reached in a given range of the rotation frequency of the wind rotor. III. CONSTRUCT DESCRIPTION The magnetic system of a multi-pole generator with an external rotor for a wind turbine is shown in Fig. 1. The external rotor of the generator consists of a outer core 1, on which the permanent magnets 2 of alternating polarity are fixed. The stator 3 of the electric generator has rectangular slots, on each of the stator teeth there is a separate winding coil, , the number of coils of each phase is 15; the number of turns of each coil - 10. The magnetic system of the electric generator with an integrated magnetic gearbox is shown in Fig. 2. The magnetic gearbox consists of an inner high-speed rotor 1, consisting of two layers of radially magnetized magnets 2, fixed segments of electrical laminated steel 4, an external low-speed rotor 5 with radially magnetized permanent magnets 6 mounted on a ferromagnetic ring core of structural steel, which is mechanically connected with a wind rotor shaft [1]",
            " For a given value of the height of the steel segments of the modulator, the dependence of the electromagnetic torques acting on the outer and inner rotors were calculated when the external rotor shifted in a given range. In this range, the maximum values of the electromagnetic torques for the external and internal rotors were determined. Then the next value of the height of the steel segments of the modulator was set, the dependence of the electromagnetic torques were calculated and the following maximum values of the torques were determined. The magnetic field picture of the generator with an external rotor is shown in the lower part of Fig.1. It should be noted that the elements of the generator magnetic circuit are not saturated. In the lower part of Fig. 2, the magnetic field of the GMG is shown, from which it is clear that the presence of a modulator leads to a complex redistribution of the magnetic field between the outer and inner rotors of the magnetic gearbox. Analysis of the results of calculations of this generator also confirmed the absence of saturation in the ferromagnetic elements of the magnetic circuit. With an increase in the height of the modulator segments, the magnitudes of the maximum torques of the outer and inner rotors increase, however, with a value of this height greater than 15 mm, the further increasing in the maximum value of torque is insignificant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001018_isef45929.2019.9097061-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001018_isef45929.2019.9097061-Figure1-1.png",
        "caption": "Fig. 1. Basic structure of SPM outer rotor spherical actuator.",
        "texts": [
            " INTRODUCTION Spherical actuators have several advantages such as a simple structure, a low moment of inertia, and a simple control method compared with conventional multi-degree-offreedom (multi-DOF) actuating systems that are composed of several single-DOF motors. Therefore, they are expected to be applied in the fields of robotics, industrial machinery, etc., and have been actively developed [1-3]. In particular, a permanent magnet synchronous spherical actuator is popular because its torque can be controlled using a torque generating equation around an arbitrary axis. For this reason, a variety of studies have focused on the permanent magnet synchronous spherical motor. As shown in Fig. 1, the outer rotor of the synchronous spherical actuator has permanent magnets, where the polarity of the permanent magnets changes in latitudinal and longitudinal direction. The stator has several coils. The rotor is driven in 3-DOF motion using currents of several phases. Therefore, the number of current phases is increased, and a control circuit becomes large and complicated. In order to solve these problems, a spherical actuator with 4 current phases has been developed [3]. However, it is pointed out that torques are decreased and there are positions where little torque can be generated around specific axis due to the decrease of the number of current phases. In this paper, in order to increase the torque, a new shape with an auxiliary pole is proposed. Furthermore, the arrangement of the coils is optimized, and a 3-DOF spherical actuator using 5 current phases is proposed. Finally, the torque characteristics of the proposed actuator are investigated using 3-D finite element analysis (FEA) and torque area method[4]. II. BASIC STRUCTURE AND TORQUE EQUATION The basic structure of our previous outer rotor spherical actuator is shown in Fig. 1[1]. Spherical shell-shaped permanent magnets are placed on the inner surface of the rotor, and their poles change by every 22.5 deg. in the latitudinal direction and every 30 deg. in the longitudinal direction. The stator has 32 coils, which are arranged by every 30 deg. in the latitudinal direction and every 45 deg. in the longitudinal direction. This actuator is driven using 16- phase currents because of two sets of 16-phase coils. The torque equation is given as follows. 1 16 1 16m m cog cog T K p K p i T p T p (1) where T is the output torque vector, Km is the magnet torque constant vector of each pole, Tcog is the cogging torque vector of each pole, i is the coil current vector, and p is the position of each coil. The current vector is a vector having the current amplitude of each phase as an element. Each torque constant vector can be obtained using the FEA. In this paper, a magnetic saturation is not considered. As shown in (1), the torque of the spherical actuator can be obtained by summing up torques generated by a simple magnetic pole model as shown in Fig. 1 (d) in each position. Therefore, in this paper, the torque in the full model is obtained by summing up torques of the simple magnetic pole model. TABLE I. SPECIFICATIONS OF SPHERICAL ACTUATOR Outer diameter of rotor [mm] 97.0 Outer diameter of stator [mm] 82.0 Residual magnetic flux density [T] 0.68 Number of coil turns 180 Movable range Around X and Y axes [deg] 40 Around Z axis [deg] 360 (continuously) Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on July 27,2020 at 02:36:03 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002113_0954407020984668-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002113_0954407020984668-Figure5-1.png",
        "caption": "Figure 5. Simplified representation of the reference theoretical model characterized by 31 parameters.",
        "texts": [
            " During the present research activity, the complete modeling of the steering, both as a separate subsystem and as part of the vehicle assembly, has always considered the assisting force, modeled as a force acting on the rack that opposes the force induced by tie rods. This force is defined by a non-linear function of the torque angle of the torsion bar. But, with the aim of mechanically characterizing the steering system, the assisting force was not modeled both in multibody models of the steering on bench and in the reduced ones such as that represented in Figure 5. In fact, this contribution is a force, applied in parallel to the system, and not a further degree of freedom and, therefore, not subjected to a possible reduction of the degree of freedoms of the system itself. The theoretical model, shown in Figure 5, has five degrees of freedom and requires the definition of 31 parameters such as inertias lengths, stiffness, damping, etc. (Table 1). Its numerical implementation was conducted in a multibody simulation environment, considered as reference for the automotive industry, MSC.ADAMS/View. The multibody model is shown in Figure 6. In Table 1 the parameters, necessary to represents the multibody model, for a total of 31 variables, are summarized. In particular, in this table the values are related to a real steering system that will be used as a test case in the paper",
            " From this observation, confirmed by other tests conducted on other examples of steering columns, the authors have defined the first hypothesis of the proposed simplified modeling procedure: the steering column inertia does not influence system dynamics and, therefore, the mass of the whole can be traced and reduced to the mass of the rack (which typically assumes values of approximately 2 kg), then assigning an unitary value to the inertia of the steering column. Considering the steering model of Figure 9(a) (repetition of that of Figure 5 but in which friction torques to the joints and the relative stiffness and rotations are represented) the equilibrium equations of the static can be written: M1 =KJ1 u1 \u00f07\u00de M2 =KJ2 u2 = M1 t12 \u00f08\u00de M3 =Kt u3 = M1 t12 t23 \u00f09\u00de with M1 applied moment to the steer, KJ1 torsional stiffness connecting the upper universal joint, u1 torsional angle associated with the steering column, M2 moment transmitted to the intermediate shaft, KJ2 torsional stiffness connecting the bottom universal joint, u2 torsional angle associated to the intermediate shaft, M3 moment transmitted to the input shaft, KT torsional stiffness representative of the torsion bar, u3 torsional angle associated with the torsion bar, and t12, t23 transmission ratios of the upper and lower universal joint, respectively",
            " This result was obtained by a series of hypotheses that was supported by a step by step comparison between the full multibody model and the simplified models realized during the research activity. In particular, the parameters that define the obtained sdof model are the mass of the rack, rack friction parameters, the equivalent stiffness of the steering column, the rack-topinion reduction ratio, a polynomial function representing the kinematic link between the rack displacement and the steering wheel rotation, all these parameters derived from geometrical and inertia characteristics of the full model (see Figure 5 and Table 1). The paper demonstrates that this reduced model is able to catch the complexity of complete multibody steering models as well as significantly speed up the calculation process in order to open the doors to realtime applications such as explicit multibody codes and simulators environment. Next step of the activity will be the definition of a methodology to obtain the sdof model starting from real steering system and from its experimental identification at bench. This could be useful if all the parameters (31 parameters) not were know"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002783_j.engfailanal.2021.105463-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002783_j.engfailanal.2021.105463-Figure16-1.png",
        "caption": "Fig. 16. DPF assembly - finite element model.",
        "texts": [
            " Gasket stresses are difficult to measure, and thus simulation models are necessary in order to study the correlation between the assembly force and the gasket compression in more detail. The gasket compression can then be correlated to the leakage as measured, to obtain a reference for future design work. A simulation model is assembled for the finite element solver software ABAQUS v. 2019 [22]. The gasket and the outer tube are modeled with solid elements to improve the mechanical contact interpretation, see Fig. 16. The 1.5 mm thick outer tube is meshed with two layers of elements through the thickness. Also solid element are utilized for the weld ring that is positioned behind the gasket and keeps the gasket in place during the assembly, Fig. 16. The inner tube with thickness 1.2 mm is modeled with shell elements to resolve bending stresses accurately. The substrate filter is omitted from the model, as it is only weakly attached to the substrate tube with an insulation mat. A general mesh size of approximately 2.5 mm results in about totally 350 000 finite elements in the model. Approximation of geometry of the FE-model is about within a distance 0.1 mm. The tube assembly load case is defined as a prescribed displacement at the inner tube attachment clips, and the assembly force is then determined as the total reaction force of these prescribed displacements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure44.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure44.6-1.png",
        "caption": "Fig. 44.6 Angular displacement curve",
        "texts": [
            " 48 g \u2022 It ca n be ar a m ax im um w ei gh t of ab ou t4 20 g \u2013 C on su m e at le as t1 /4 fo ot pr in to f th e pl at e \u2013 C on su m es th e bo tto m sp ac e of pl at e as w el lh en ce cr ea te hi nd ra nc e w hi le ho ld in g th e pl at e \u2013 C an be ar a m ax im um w ei gh to f on ly 22 0 g \u2013 It ha s a he av yw ei gh tc om pa re d to al lo th er de si gn s an d to ho ld it on a of pl at e re qu ir es m or e fo rc e by us er du e to its m om en tu m \u2013 C an no ta cc om m od at e di ff er en tv ar ie ty of gl as se s an d cu ps \u2013 T hi s at ta ch m en tw as no ta bl e to ho ld th e cu ps an d flu te gl as se s \u2013 C la m p de si gn is no tfi tf or m an y pl at es de si gn \u2013 Pi vo tp oi nt s ar e lo w er th an th e C G of gl as s w hi ch le ad s to m al fu nc tio ni ng of gi m ba lm ec ha ni sm an d ul tim at el y tu m bl in g of gl as s \u2013 O nl y su ita bl e fo r ra ng e ad op te d co ns id er in g In di an m ea l \u2022 D es cr ip tio n \u2013 D ra w ba ck s Its open-ended base provides a spring action that allows a number of different shapes and sizes to fit into it. Also due to the spring action, it provides a good grip to the glass and cup which helps in stabilizing the glass or cup, so the drink will not spill out of the glass or the cup. In Fig. 44.6, point \u2018O\u2019 is the hinge point where basket and clamp are attached and due to contact, there is a frictional force that acts opposite to the angular velocity. Point \u2018C\u2019 is the center of gravity (COG) of the basket and the glass which is filled with some liquid. If any slight force is applied to the basket or glass, the COG will shift at an angle \u2018\u03b8 \u2019 from its mean position. Here: FG = Force of gravity, FA = force of gravity along the line of action, FP = force of gravity perpendicular to the line of action, and F t = Tension force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure4-1.png",
        "caption": "Fig. 4. Commutator and windows over an orbit revolution of the rotor.",
        "texts": [
            " 3(a), as dictated by the commutator. Fig. 3(c) represents the view of windowsmanifold when viewed from the rotor side. Each one of the inner slots represented in Fig. 3(c) is connected to each of the TSV in the rotor-roller set. The commutator and rotor are connected by the floating spline joint (Fig. 2) and thus they move synchronously. For the continuous operation of the orbit motor, high pressure fluid is provided at a different angle to the TSV, as represented by the cross-sectional view of Fig. 3(b). Fig. 4 represents the windows under consideration along with the commutator (colored in gray) over an orbit rotation. The commutator performs the orbiting action, subjecting each of the windows in the windows manifold, 5 (Fig. 2), to the inside or outside of itself. From Fig. 4, it can be observed that the windows are exposed to the inside of the commutator or the outside which correspond to the outlet (colored in blue) and inlet environments (colored in red) respectively. For a certain duration of angle, the window area is completely enveloped by the commutator thus providing zero cross-porting. The structure of the developed simulation model is presented in Fig. 5. The four modules of the simulation model are: the geometric pre-processor module, the mechanical pre-processor module, the gap module and the fluid dynamics & force module",
            " All these quantities, essential to evaluate both the fluid and the force features are provided as a lookup table with respect to the rotation angle of the orbit motor. The commutator geometric module evaluates only the porting characteristics realized by the commutator with respect to the windows manifold (Fig. 2). As explained earlier, the commutator separates the high-pressure and the low-pressure environments thus subjecting each of the windows to high and low pressures sequentially while the orbit motion is being performed (Fig. 4). The commutator geometric module evaluates the instantaneous area of communication and hydraulic diameter of each of the windows using GSL libraries. Fig. 10 represents part of the output generated by the commutator geometric module. For one of the windows, the instantaneous port communication area to the inlet and outlet environments is presented for one orbit revolution of the rotor. It can be noted that the inlet and outlet areas do not intersect i.e., zero crossport. This construction leads to the TSVs connecting to either of the ports or none but not both instantaneously"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002423_012128-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002423_012128-Figure1-1.png",
        "caption": "Figure 1. Re-entrant cubic auxetic structure.",
        "texts": [
            " Designing an auxetic structure that has a simple geometry with a unit cell that provides room to be easily scaled without compromising its auxetic nature is a challenging task. This paper has addressed this my modifying the idealised unit cell derived from an auxetic foam [5] to make it more suitable for mass production. The auxetic structure is selected and modified based on four key factors. The structure should have less repeating units, better scalability, use standardised raw materials and mainly the feasibility for mass production. A new re-entrant cubic auxetic structure Figure 1 is developed based on these key factors. A repeating unit shown in Figure 2 is a sub unit that constitute a unit cell (Figure 3). This unit cell exhibits auxetic properties. The lesser the repeating units the easier would be the assembly of the unit cell. These unit cells must have the provision to link with other such unit cells so it can be scaled with minimal effort. The materials used to fabricate the auxetic structure must be available in standard dimensions. This eliminates the additional step of in house fabrication. Finally the shape of the structure must be simple and cost effective to be assembled. The cubic unit cell (Figure 1) that resembles a systematically collapsed cube is re-entrant in nature. This cell can be divided into repeating units that can be further subdivided into standard geometric shapes. Based on the availability of standard sized raw materials aluminium cubes (6.1 mm) and steel links (\u00d8 1.3 mm) are selected. These are fixed parameters in the unit cell. The repeating unit (Figure 2) consists of 5 steel links assembled in a particular formation to facilitate assembly with other repeating units. The re-entrant cubic unit is an assembly of 6 repeating units (Figure 4)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000129_978-3-319-64301-4_20-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000129_978-3-319-64301-4_20-Figure11-1.png",
        "caption": "Fig. 11 Mobile interface of the adwisar system, showing the assistance provided for a step in the work process",
        "texts": [
            " The main screen below shows two documents the system determined to be relevant to the employee in the current situation (the bill of materials for the current and a standard operating instruction for the current work order). The worker can click on the documents to open them. The other relevant tab is on the top right, \u201cAssistenz\u201d meaning \u201cassistance\u201d (see Fig. 10). There, the system displays work procedures that are applicable in the current context. If the employee selects one of the work procedures, she will see instructions for each step of the process. Figure 11 shows the interface when the user has entered a work procedure. The system displays detailed information on the precise action to perform. The action of the worker triggers a magnetic detection sensor attached to the frame, and the sensor sends the current force value to the Kerlink data management system. In this example, the raw sensor data is translated by Kerlink into side panel2PlacedInRHGroove, meaning that the operator inserted side panel 2 into the right-hand groove correctly. The Kerlin system then sends this message to the adwisar machine information service through a post request"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001860_iecon43393.2020.9254702-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001860_iecon43393.2020.9254702-Figure2-1.png",
        "caption": "Fig. 2. Specifications of the dual-arm system.",
        "texts": [
            " To the best of the authors\u2019 knowledge, this is the first work where the NMPC-based trajectory generation with redundancy resolution strategy is implemented for the dual-arm aerial manipulator with all computations done online and onboard. This section describes the mechanical construction of the dual-arm aerial manipulator utilized to validate the presented approach in Section III. The aerial robot is composed of the custom-build hexarotor and dual-arm sytem, consisting of the dual-arm base and two human-size arms as shown in Fig. 2. Each arm provides six DOFs for the end-effector positioning with the following joint types: shoulder prismatic, shoulder yaw, shoulder pitch, elbow pitch, wrist pitch, and wrist roll. A DC motor with an encoder from DFRobot is used for the shoulder prismatic joints, while Dynamixel X-series actuators are utilized for the rest of the joints. The total mass of the dual-arm system is 2.5 kg with the maximum payload per arm of 1.0 kg. The prismatic joints are employed to achieve a better flight performance by dynamically adjusting the dualarm COG as well as to increase the workspace and reach of the arms",
            " Hence, attaching multi-DOFs robotic arms to multirotor UAVs leads to the whole new systems with exceptional redundancy. Then, it is worth to specify the complete state of the dual-arm aerial manipulator which includes the position (rb = [xb yb zb] T \u2208 R 3) and UAV orientation represented by the Euler angles (\u03a6b = [\u03c6 \u03b8 \u03c8]T \u2208 R 3) with respect to the world-fixed reference frame, and the dual-arm joint position vectors qm = [qm1 , . . . , q m 6 ]T \u2208 R 6, with m = {1, 2} for the right and left arms, respectively (see Fig. 2). Thus, the complete state of the dual-arm aerial manipulator is defined as follows: \u03c3 = [rTb \u03a6 T b q1,T q2,T ]T \u2208 R 18 (1) However, having such high-dimensional state increases the computational burden, which, in its turn, can hinder the realtime implementation of the proposed trajectory generation algorithm, considering the available onboard computational power. Hence, it was decided to exclude q 1,2 1 and q 1,2 6 from consideration in the proposed framework, in order to decrease the computational load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001701_0954406220964512-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001701_0954406220964512-Figure1-1.png",
        "caption": "Figure 1. A 4-DOF planar manipulator with two obstacles in its workspace.",
        "texts": [
            " By inserting equation (5) into equation (3), the overall inverse kinematic solution with task priority is given _h \u00bc J#1 _r1 \u00fe J\u0302 # 2 _r2\u2013J2J # 1 _r1 \u00fe I\u2013J#1 J1 n o \u00f0I\u2013J\u0302# 2 J\u03022\u00dez (6) where the Hermitian and idempotent8 property of the projection operator is used for simplification. Similar analogy has been drawn at the acceleration level to minimize the joint torques locally.33 For the second manipulation variable, we can have desired trajectory or a criterion function like potential function. In the case of r2 \u00bc h, equation (6) can be simplified using J2 \u00bc In as follows, _h \u00bc J#1 _r1 \u00fe \u00f0I\u2013J#1 J1\u00de_r2 (7) where; \u00f0I\u2013J#1 J1\u00de# \u00bc \u00f0I\u2013J#1 J1\u00de and the idempotency of \u00f0I\u2013J#1 J1\u00deare used. Obstacle avoidance using potential function Figure 1 shows, the physical model of a planar 4-DOF manipulator with two rectangular shapedobstacles placed in the workspace, of dimension \u00f0LOi \u00f0m\u00de HOi \u00f0m\u00de\u00de: The centre of gravity of the two is marked as (CGxi, CGyi). The point pi shows the mid-point of the ith link of the manipulator which is used to keep the links away from the obstacle while tracking the trajectory. The dynamics of a general n DOF manipulator, is represented as s \u00bc M h\u00f0 \u00de\u20ach \u00fe c h; _h \u00fe g\u00f0h\u00de (8) For planar manipulators, the effect of gravitational forces can be neglected, so equation (8) can be modified as: s \u00bc M h\u00f0 \u00de\u20ach \u00fe c h; _h (9) Let the first manipulation variables is as follows: r1 \u00bc xtip ytip (10) where r1 represents the position of the end-effector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002315_j.matpr.2020.12.487-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002315_j.matpr.2020.12.487-Figure6-1.png",
        "caption": "Fig. 6. Analysis results of insole (a) dorsiflexion; (b) plantar flexion; (c) cylinder support at mid foot; (d) cylinder supported at fore foot and hind foot.",
        "texts": [
            " Material Contact Analysis Results Start and end Gait cycle With cylinder supports Dorsiflexion Plantar Flexion At mid foot At hind and rare foot 1. Ninja Flex Von Mises 0.26176 0.40548 955.6 326.1 Deformation 0.0090138 0.017499 212.98 21.479 2. Flex PLA Von Mises 0.26135 0.40493 857.12 238.22 Deformation 0.00090266 0.0017601 32.261 1.6602 3. PLA Von Mises 0.2629 0.40822 797.28 269.6 Deformation 0.00039878 0.00075819 18.956 3.0826 tion. Loads and boundary conditions are applied at heel, metatarsal and all degrees of freedom on bottom face. Results of Von Mises stresses and deformation are shown in Fig. 6. Results are shown in Table 3. Lattice structures are most preferred for mass and topology optimization. BCC, OCTO, IWP lattice structures are modelled in 32 32 16 mm space using CATIA software. In BCC lattice linear cellular structure, struts are modelled with circular cross section of diameter 0.8 mm in 4 mm cuboid space (A). With the similar sizes, BCC lattice curved struts are developed (B). Pattern (B) is modified by placing 0.4 mm plates at 4 mm spacing (C) and 2 mm spacing (D). OCTO surface based unit cell is modelled with hollow sphere at face centred (E)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000715_s00419-020-01687-2-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000715_s00419-020-01687-2-Figure1-1.png",
        "caption": "Fig. 1 a Image of the Szabad(ka) II hexapod walker robot; b image of the robot\u2019s SolidWorks model",
        "texts": [
            " As a result of the optimization, the authors determined the parameters of the friction, backlash and self-locking effects in the mechanics of the robot leg. The validation was performed by using the joint positions and motor currents measured on the physical robot. In the fourth chapter, the Conclusion, the achieved results are summarized and the conclusions are drawn. The accuracy of the validated model has increased significantly by expanding the model with a detailed description of friction, self-locking of the reductors and gear backlash. Szabad(ka) II is a hexapod walker robot with three degrees of freedom (DOF) per leg (Fig. 1a). The structure is powered by 18 DC servo motors fitted with encoders, manufactured by Faulhaber. The motors are driven with self-designed motor drivers. Most of the structure is made of aluminum. Torque transmission in the joints is achieved by reductors, fitted with bevel gears, as well as with plain bearings (manufactured by IGUS) and ball bearings (manufactured by SKF). In the center of the robot\u2019s body are the electronic elements and the battery. Each leg consists of three segments: tibia, femur and coxa. The coxa\u2013thorax (Link 1) joint can move along the horizontal plane, while the coxa\u2013femur (Link 2) and the femur\u2013tibia (Link 3) joints can move along the vertical plane. The arrangement of the legs and leg segments can be seen in Fig. 1b. The robot model consists of several nested units. These units can be classified into the following major categories according to their characteristics: \u2013 Kinematics and dynamics, \u2013 Control, electronics and drive, \u2013 Non-ideal mechanical phenomena. The dynamic motion equation of the robot is widely known and can be generally expressed by Eq. (1): \u03c4 = H(q)q\u0308 + C(q, q\u0307)q\u0307 + g(q) + J (q)T fext (1) where in the case of n joints (in case of our robot n = 18+6 = 24), \u03c4 is the n\u00d71-dimensionalmatrix of the driving torques/forces, q is the n \u00d7 1-dimensional column vector of the wrist coordinates, q = (q1, q2, ",
            " The ports for connecting the robot\u2019s legs (1RF-6LR) also use rigid transforms to connect to the other elements of the body. A seventh port (F) can be used to connect the body and the ground. The body can move completely freely in relation to the ground. The 6-DOF joint connecting the body and the ground has the role of determining the body\u2019s position and orientation. The force interactions between the robot and the ground arise only through the feet, when a particular leg has ground contact. The leg sub-systems attached to the body are named in relation to the position of the legs, as shown in Fig. 1b: RF\u2014right front, LF\u2014left front, RM\u2014right middle, LM\u2014left middle, RR\u2014right rear and LR\u2014left rear. The legs consist of further sub-systems and the joints connecting them (Fig. 3). These sub-systems represent the three leg segments (tibia, femur and coxa), whereas the revolute joints correspond to the three joints (Link 1, Link 2 and Link 3). The three sub-systems consist of solids corresponding to the components and rigid transforms that connect them and determine orientation. Each joint has been fitted with an actuation (torque) and two sensing (position and velocity) ports"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000842_j.asr.2020.03.021-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000842_j.asr.2020.03.021-Figure5-1.png",
        "caption": "Fig. 5. Optimal cost J for moment control.",
        "texts": [
            " These attributes of the dynamic-programming-based approach can help in determining the parameters of the quadratic cost function in order to reach desired steady-state errors. For instance, the parameters in Table 5 are designed such that the maximum width of the dead-zones for position and attitude are less than 5 mm and 0.6 degrees, respectively. Furthermore, as a result of implementing Eq. (18), the optimal cost for attitude control reaches zero at multiples of 2p as can be seen in the example shown in Fig. 5. For each of the state variables, 700 logarithmically spaced points are generated within the specified ranges. Logarithmic spacing effectively increases the resolution around the origin - where it is most desired - by creating a denser meshing, while eliminating unnecessary computations in peripheral regions. For the control variables, 41 linearly spaced points are generated to include all the feasible actions that the thrusters are capable of producing. Furthermore, the terminal cost (first term of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001354_s10015-020-00628-0-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001354_s10015-020-00628-0-Figure1-1.png",
        "caption": "Fig. 1 The collective foraging task with poison",
        "texts": [
            ", , 1 3 The objective of the task is that the robots should distinguish between the two kinds of objects, namely food and poison, and transport only food objects to the nest. A single robot cannot distinguish between food and poison objects by itself, i.e., the robotic swarm has to perform collective cognition to distinguish between the two objects. Additionally, food and poison objects are too heavy for a single robot to move; therefore, robots have to cooperate with each other to accomplish the task. The experiments are carried out in computer simulations using the Box2D physics engine1. The simulation environment has a regular octagon field, as shown in Fig.\u00a01. A circle-shaped nest with a radius of 15\u00a0m is located in the center of the field. At the beginning of the simulations, robots are positioned inside the nest. The initial positions of the food and poison objects are randomly determined. There are always five food objects and five poison objects in the field. A new food or poison object will be generated with a random position when it is transported to the nest. The radii of food and poison objects are set to 5.0\u00a0m and 2.5\u00a0m, respectively. The food and poison objects are set to have the same weight; in particular, at least four robots are required to move an object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003025_tee.23411-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003025_tee.23411-Figure8-1.png",
        "caption": "Fig. 8. IPMSM. (a): Rotor; (b): Installation diagram. (1) motor shaft; (2) permanent magnet; (3) rotor lamination; (4) stator winding; (5) IPMSM body",
        "texts": [
            " XT2 e + YTe + Z = 0 (26a) X = \u2212 L + Rc(b2 1 + b2 2 ) L 3 (26b) Y = [(a1b1 + a2b2)\u03d5f \u2212 (b1c1 + b2c2) L] \u00b7 Rc(\u03d5f + Lid ) (26c) Z = [Rs id + Rc(a 2 1 + b2 1 )id + Rc(a1c1 + a2c2)] \u00b7 3(\u03d5f + Lid )3 (26d) where \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 a1 = 0.5\u03c9eLd Lq Para, a2 = Ld RcPara b1 = \u2212RcLq Para, b2 = a1 c1 = 0.5\u03c9e\u03d5f Lq Para, c2 = Rc\u03d5f Para Para = 2\u03c9e 4R2 c +\u03c92 e Ld Lq The calculated power loss curves of three different strategies (1 Nm, 0\u20134 krpm) are drawn in Fig. 7. As can be seen, MEA has the lowest power loss comparing with CLMC and MTPA. 4. Experimental Results The experiments were implemented on the experimental platform with an IPMSM. The structure of the IPMSM is shown in Fig. 8. Additionally, the experimental inductances of the IPMSM are presented in Fig. 9, and the fitted inductances are defined as:{ L\u0302d = fd (is ) L\u0302q = fq (is ) (27) Besides, the whole experimental platform is presented in Fig. 10. The parameters of the experiment system are shown in Table I. The VSI module is an IPM (PM50RLA060), and the DSP2812 is chosen as the control processer, which is widely applied for the IPMSM drive system. Additionally, the simplified schematic of the platform is presented in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001378_s11668-020-00961-3-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001378_s11668-020-00961-3-Figure6-1.png",
        "caption": "Fig. 6 Assembly structure of the middle of the spring",
        "texts": [
            " Table 2 Working conditions of the composite leaf spring Stiffness matching Free stiffness Clamping stiffness Load condition Vertical Steering Braking Fatigue U-bolt pre-tightening Hook Fig. 4 Static stiffness of the composite leaf spring and the steel leaf spring Stress Analysis The longitudinal composite leaf spring is subject to the vertical, lateral, and longitudinal forces acting on the automobile chassis. The force acts on each part of the leaf spring differently. Clamping regions of U-bolts are mainly affected by bolt pre-tightening forces. Figure 6 is the clamping structure diagram of the U-bolt in the middle of the leaf spring. The structure is composed of the U-bolt, the spring body, the cover plate, the up plate and the down plate. The spring body is made of composite materials, while other parts are metal. Different U-bolts are selected according to different vehicle loads. Bolts with different grades and diameters have different torques. Standard bolt torques are shown in Table 3. Figure 7 shows the stress distribution of various parts in the leaf spring after U-bolt pre-tightening"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002407_ieeeconf35879.2020.9329622-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002407_ieeeconf35879.2020.9329622-Figure2-1.png",
        "caption": "Fig. 2. Miura-Ori unit cell, where a = 24 mm, b = 24, mm, \u03b3 = 65 .\u030a Equations from [7]",
        "texts": [
            "55 mm diameter coax cables are referenced together to prevent any unwanted common-mode resonance [6]. III. IMPLEMENTING PERIODIC ARRAYS ON MIURA When designing a classic periodic array, a typical unit cell is on a square or rectangular lattice. Alternatively, Miura unit cells are defined using a parallelogram of sides a and b and an acute angle gamma (\u03b3). The formation of the different states of the Miura pattern are controlled by the fold angle theta (\u03b8), referenced off the XY plane [7]. The unit cell dimension [7] can be expressed in terms of length L, width S, and height H, as given in Fig. 2. Therefore, the total volume of the unit cell is (2L+V)\u00d72S\u00d7H. This work was supported by the Air Force Office of Scientific Research (AFSOR) grant FA9550-18-1-0191. 1685978-1-7281-6670-4/20/$31.00 \u00a92020 IEEE APS 2020 20 20 IE EE In te rn at io na l S ym po siu m o n An te nn as a nd P ro pa ga tio n an d N or th A m er ic an R ad io S ci en ce M ee tin g | 97 8- 1- 72 81 -6 67 0- 4/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D O I: 10 .1 Authorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 16:56:20 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003284_s40430-021-03157-4-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003284_s40430-021-03157-4-Figure3-1.png",
        "caption": "Fig. 3 Virtual representation of the Delta robot, highlighting the corresponding geometric parameters",
        "texts": [
            " The Delta kinematic structure is a parallel mechanism with only 3 translational DoFs at the end-effector, which is always parallel to the fixed base of the robot. The 3 rotational DoFs are constrained by the parallelogram shape [50] of the arms that connect the end-effector to the actuated links, as shown in Fig.\u00a01. This means that both connecting rods have the same length, and both extremities are separated by an equal distance. Each of these rod ends are connected to the neighboring links through spherical joints. The geometric and inertial parameters of the Delta robot are listed below5 and indicated in Fig.\u00a03. \u2022 Base radius: r = 0.4\u00a0m; \u2022 Arm (in red) length: a = 0.4\u00a0m; \u2022 Parallelogram (in black) length: b = 1.0\u00a0m; \u2022 End-effector radius: h = 0.15\u00a0m; \u2022 Arm mass: ma = 0.4\u00a0kg; \u2022 Parallelogram mass (each bar): mb = 0.1\u00a0kg; \u2022 End-effector mass: mh = 0.2\u00a0kg. As Fig.\u00a0 3 suggests, this Delta robot is symmetrical, meaning that each arm is equally set apart from the other two. This joint placement is parameterized in polar coordinates around the z axis with i angles, being 1 = 0 (the first arm is the one aligned with the x axis), 2 = 120\u25e6 and 3 = \u2212120\u25e6 . Following this order, joint motion angles are denoted as i , with positive values referring to an upward tilt of the corresponding arm. All these parameters enable building the required mathematical representations of the robot, described in detail in the following subsections",
            " The Delta robot always has two possible solutions for this problem [43]. One solution corresponds to the end-effector below the fixed base, which is physically consistent with the mechanical structure of the robot (see Figs.\u00a01 and 3). Contrarily, the second solution relates to the end-effector above the base of the robot, being of no practical interest. This poses the criterion for choosing the right solution: a negative value for the z coordinate, considering the base frame orientation shown in Fig.\u00a03. To develop the forward kinematics equations, dual quaternion algebra [16], an extension of quaternions [13, 14], will be applied. A quaternion q is represented in the following form: where , x, y, z \u2208 \u211d , and i,\u00a0j,\u00a0k are the three imaginary components with the following properties: A quaternion can also be represented as a pair (s, ) , where s \u2208 \u211d is the scalar part, and \u2208 \u211d 3 , the vector part. The correspondence between the two notations is given as follows: Dual quaternions are introduced by joining the concepts of quaternions and dual numbers, establishing an eightdimensional manifold",
            " This structure undergoes final training on a dataset of 1,000,000 (one million) samples, further reducing the root mean square error of the estimation to 0.4472\u00a0mm. The computational efficiency of this algorithm is also evaluated by computing the forward kinematics of a set of 1000 (one thousand) samples, using both the trustregion-dogleg [55] and the trained neural network. While the numeric algorithm completes the task in 4.174\u00a0s, the neural network takes only 0.025\u00a0s, representing a speedup of 167 times. Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:440 1 3 440 Page 8 of 11 The physical system shown in Fig.\u00a03, the control loop of Fig.\u00a02, and the required computations were implemented inside a Simulink6 environment, with the aid of Simscape Multibody and the Deep Learning Toolbox. The impedance control law (Eq.\u00a02) is first loaded with K = 50\u00a0N/m and B = 10\u00a0Ns/m for all three axes. The chronological sequence of events is listed below. \u2022 The robot starts with all three actuated arms in horizontal position, which corresponds to the end-effector at x = y = 0 and z = \u22120.76\u00a0m. Reference position is set at x = y = 0 and z = \u22121\u00a0m (the analytic solution of the forward kinematics problem would fail at these points due to numeric singularities); \u2022 At t = 5\u00a0s, a 5\u00a0N, 3\u00a0rad/s perturbation starts at the x axis and persists until t = 10\u00a0s; \u2022 From t = 15\u00a0s to t = 20\u00a0s, the y axis is disturbed with the same parameters used for the previous event; \u2022 At t = 25\u00a0s, the z axis experiences a step disturbance of 5\u00a0N in the negative direction which holds until t = 30\u00a0s; \u2022 Last, reference position shifts 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000380_0954406219897688-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000380_0954406219897688-Figure6-1.png",
        "caption": "Figure 6. Bearing experimental setup from CWRU.",
        "texts": [
            " The SNRf of raw signal, VM2RLbased denoised signal and VMD-based mode were calculated as 9.6021 dB, 9.0744 dB and 9.4008 dB, respectively. The proposed method shows the best denoising effect in frequency detection and time-frequency analysis. These further indicate that the proposed method is effective in revealing the transient characteristics reflecting fault information. Case I: The vibration data from CWRU The data acquired by an experimental platform from CWRU Bearing Data Center29 were firstly analyzed. The experimental system is illustrated in Figure 6, where accelerometers are installed to the motor\u2019s case with magnetic bases, at a sampling frequency of 12 kHz. Single point faults with inner race, rollingelement and outer race defects were set on the testing drive-end bearings (deep groove ball bearing with the Type 6205-2RS JEM SKF) separately at bearing elements using electric discharge machining method. As an example of multi-modulation bands representing in a type of fault, rolling-element vibration data were used as representative in this case"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure22-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure22-1.png",
        "caption": "Fig. 22 Three-dimensional graph",
        "texts": [
            " In this mechanism, a geometric constraint is added at kinematic joint C, and the upper nipper 4 which combined with the lower nipper 2 to form one component is gradually moved from position b to position a. Based on the principle of augmented Assur groups, the metamorphic nipper swing mechanism is divided into a fixed-axis rotating active part, an Assur group RPR and an augmented Assur group RP\u2013RR\u2013RR\u2013R, as shown in Fig.\u00a021. According to geometric and inertia properties of the metamorphic nipper swing mechanism in Table\u00a0 3, a three-dimensional model is established in SolidWorks, as shown in Fig.\u00a022. The initial distance between slider 5 and H is 172.37\u00a0mm, and the initial position of the component 9 is 3 \u22155\u00a0rad. Assuming that the active part 9 rotates at a constant speed of 6 r/min (motion period is 10\u00a0s), the dynamic simulation is carried out in SolidWorks virtual prototype environment, and the relationship between the driving torque and time of the metamorphic nipper swing mechanism is obtained. 1 3 When the metamorphic nipper swing mechanism is in gradually closed and gradually opened configurations, the dynamic equations can be obtained by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002605_s11071-021-06365-8-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002605_s11071-021-06365-8-Figure1-1.png",
        "caption": "Fig. 1 Illustration of the robotic penguin. a Conceptual design; b robotic prototype",
        "texts": [
            " Ultimately, the conclusion and future work are presented in Sect. 7. Notations : J f (qd) is the Jacobian matrices of function f with respect to qd . \u2016\u03be\u20162Q stands for Qweighted norm which is defined as \u03beT Q\u03be . Specially, all weight matrices (e.g., Q, P , and R) are diagonal positive-definite matrix. In addition, \u03be(k + i |k) means the prediction states of k + i time step at k time step, and \u03be(k+ i) denotes the actual states at k+ i time step. 2.1 Robotic penguin modeling The robotic penguin propelled by MPF is illustrated in Fig. 1, in which the total length of the body and the total mass is 0.75 m, and 19.4 kg, respectively. The corresponding specifications of the robotic penguin are provided in Table 1. Particularly, the pectoral fins have 2-DOF movement capability including flapping and feathering motion. The flapping motion mainly generates power through the reversed Kaman vortex street, while the feathering motion mostly controls the yaw direction. Owing to the movement of pectoral fins is not suitable for flexible 3D motion, a buoyancy-driven system composed of a water injector and a piston is equipped to realize ascending and diving motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000882_j.cagd.2020.101870-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000882_j.cagd.2020.101870-Figure12-1.png",
        "caption": "Fig. 12. Angle notation in Burmester\u2019s mechanism.",
        "texts": [
            " For construction of the remaining vertices and for details and references, see Wunderlich (1968). We will show that Dixon\u2019s angle condition is necessary for the movability of the Burmester mechanism in a special case. Theorem 3.3. If the linkage in Fig. 11, left, is flexible and the two lower small quadrilaterals satisfy a \u00b1 b \u00b1 c \u00b1 d = 0, then the angles \u03c91 and \u03c92 remain either equal or complement each other modulo \u03c0 during the motion. Proof. Consider the lower half of Burmester\u2019s mechanism and denote some of its angles as shown in Fig. 12. In the variables u = cot \u03d51 2 , v = cot \u03c81 2 = tan \u03d52 2 , w = cot \u03c82 2 the configuration space of the coupled quadrilaterals is described by the system P (u, v) = 0, Q (v, w) = 0, with both equations of the form (4). In other words, the complexified configuration space is the so called fiber product C12 = {(x1, x2) \u2208 C1 \u00d7 C2 | f1(x1) = f2(x2)} of the configuration spaces C1 and C2 of the two quadrilaterals. C12 C1 f1 C2 f2 u \u2208CP1 v \u2208CP1 w \u2208CP1 If the branch sets of the maps f1 and f2 on the diagram don\u2019t coincide, then C12 is a hyperelliptic curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001943_012024-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001943_012024-Figure1-1.png",
        "caption": "Figure 1. Forestry modular tillage implements: (\u0430) front-mounted drum chopper; (b) rear-mounted single-row harrow; (\u0441) rear-mounted two-row harrow.",
        "texts": [
            " These are wheels, suspension elements and attachment mechanisms for working equipment. Also, additionally, a spherical element was installed to compensate for the mass-inertial characteristics of the non-simulated stationary elements of the tractor. He has the ability to change the spatial position and mass. By adjusting these parameters, the mass-inertial characteristics of the tractor model were brought to the parameters of a real tractor. The tractor model was equipped with front and rear mounted forestry modular tillage implements [4] (figure 1). The following implements were used in the modeling: front-mounted drum chopper (figure 1a), rear-mounted single-row harrow (figure 1b); rear-mounted two-row harrow (figure 1c). Figure 2a shows three typical unit layout options: unit with a single-row rear-mounted implement (scheme Forestry 2020 IOP Conf. Series: Earth and Environmental Science 595 (2020) 012024 IOP Publishing doi:10.1088/1755-1315/595/1/012024 0 + 1), two-row rear-mounted implement (scheme 0 + 2), and with front-mounted and two-row rear-mounted implements (scheme 1 + 2). Depending on the applied reforestation technology and operating conditions, the implement s can be equipped with sets of various working tools"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000311_s12541-019-00228-4-Figure24-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000311_s12541-019-00228-4-Figure24-1.png",
        "caption": "Fig. 24 Stress analysis of the lower panel",
        "texts": [
            " The EBDOP-1 and EBDOP-2 designs show an 1 3 ~ 890\u00a0MPa increase at the attachment point, with additional hotpots and a larger area of the lower panel with stresses > 350\u00a0MPa. The software-based optimised design (SBDOP) shows a 224\u00a0MPa increase in stress at the right-hand side body attachment compared to the baseline but a reduction in the number of hotpots above 350\u00a0MPa. The EBDOP-1 and EBDOP-2 designs both show an ~ 270\u00a0MPa increase at the attachment point. The stress analysis of the lower panel is shown in Fig.\u00a024 and presented in Table\u00a06. Software-based optimisation method (SBOP) is better method to reduce stress lower panel. The software-based optimised design (SBDOP) shows very similar stresses to the baseline for the left-hand side tower. The EBDOP-1 and EBDOP-2 designs both show an ~ 336\u00a0MPa decrease at the outermost attachment points. This is the result of adopting a single pressing, which provides better local stiffness. However, additional hotpots above 350\u00a0MPa that do not appear on the baseline are observed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure4-1.png",
        "caption": "Fig. 4. Inverse kinematic singularity.",
        "texts": [
            " Based on the Jacobian matrices in Sec. 3.1 [32], the kinematic singularity of the 2UPR-PRU PKM can be classified as inverse, forward, and combined types, which are discussed in the following sections. 3.2.1 Inverse kinematic singularity When =qJ 0 and \u2260\u03b7J 0, the PKM is in the inverse kine- matic singular configuration. =qJ 0 occurs when one of 11q J , 22q J , or 33q J is equal to zero. From Eq. (10), one can derive that the 2UPR-PRU PKM is in the inverse kinematic singularity only if limb 3 is perpendicular to slider 3OB . Fig. 4 shows the configuration. 3.2.2 Forward kinematic singularity When =\u03b7J 0 and \u2260qJ 0, the PKM is in the forward kinematic singular configuration. Eq. (10) allows for the derivation that the 2UPR-PRU PKM is in the forward kinematic singularity only when limb 1 or 2 is co-linear with 1 2A A . Therefore, two forward kinematic singular configurations occur, as shown in Fig. 5. 3.2.3 Combined kinematic singularity Only when =qJ 0 and =\u03b7J 0 are satisfied simultaneously, the PKM is in the combined kinematic singular configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001431_tec.2020.3017077-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001431_tec.2020.3017077-Figure8-1.png",
        "caption": "Fig. 8. 3-D FEM and the mesh result of the axial flux disk motor.",
        "texts": [
            " Then Maxwell's equations t      D H J (6) t t           B H E (7) can be written in terms of the current vector potential T and the scalar magnetic potential \u03c9\uff0cnamely T-\u03c9 formulation 2 =0 ( ) ( )J t                T T (8) After obtaining the governing equations, various field quantities in the solution domain can be calculated by adopting the finite element numerical method [13, 14]. According to Table I, II and the measured material electromagnetic properties, the software ANSYS Maxwell is used to establish the 3D FEM of three-phase asynchronous disk motor in RT environment. Fig. 8 shows the 3D FEM, the stator, the rotor and 3D finite element mesh result of three-phase asynchronous disk motor. By setting different geometric parameters and working parameters in 3D FEM, the operating performances and electromagnetic force behaviors of the disk motor can be obtained. In order to verify the 3D FEM established for RT environment, the experimental setup is constructed to test the toque and electromagnetic force characteristics of disk motor, as shown in Fig. 9. The setup mainly consists of the disk motor (stator and rotor), the variable frequency controller (ABB ACS355), the electric eddy current brake (providing an adjustable load torque) and load controller, the speed/torque sensor and the digital oscilloscope, the spoke-structure pull-pressure sensor (\u00b12000 N, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001716_ccs49175.2020.9231465-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001716_ccs49175.2020.9231465-Figure3-1.png",
        "caption": "Fig. 3. The hierarchical structure of the custom vision-based tactile sensor.",
        "texts": [
            " The camera is fixed on the framework through bolts and nuts, and uses USB connection to transmit images or videos to PC. Although the camera stream of BRIO Webcam can achieve 1080P at 30fps, it has a large length of 100mm. When there is miniaturization or other requirements, the camera can be replaced with a common camera with fisheye lens, as is used in the prototype FingerVision, with a camera stream of only 320x240 at 30- 60fps. The most important component of the sensor is a specific silicone elastomer with a multi-layer structure, as shown in Fig. 3. From the outermost to the innermost layer, they are: hard silicone shell with markers array inside, soft silicone pad and hard acrylic layer. Different layers of the silicone elastomer are naturally stuck together during curing process. The entire elastomer is supported by the innermost acrylic layer with the soft silicone pad fixed by bolts. Compared with the silicone pad used in the prototype with hardness of 16HA, that of the custom sensor is less than 5HA. Hence, a hard silicone shell coats the soft silicone pad to protect the elastomer just like epidermis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002520_iceta51985.2020.9379240-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002520_iceta51985.2020.9379240-Figure8-1.png",
        "caption": "Fig. 8. Model of RC-car equipped with Raspberry Pi minicomputer, ultrasound, infrared and light sensors",
        "texts": [
            " Downloaded on May 17,2021 at 10:01:46 UTC from IEEE Xplore. Restrictions apply. problems. The model of the hydraulic system with three connected tanks (Fig. 7), which is also built on the Arduino platform, stands out from the presented models in terms of size and construction. We put it in the pocket lab series with regard to its mobility. Although it is a model filled by a liquid, this system can be carried in a case in any position, as it is equipped with a shut-off valve. The model of the RC-car (Fig. 8) is a little bit different from the others, because it does not use the Arduino platform, but is built on the Raspberry Pi minicomputer. This laboratory model allows you to perform low-level control actions, depending on the installed sensors of distance, light or infrared signals. But, thanks to the power of the minicomputer, also more complex control tasks that require image recognition and artificial intelligence may be accomplished. The size of the Raspberry Pi minicomputer is still comparable to a credit card that allow small dimensions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002279_1.4006737-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002279_1.4006737-Figure4-1.png",
        "caption": "Fig. 4 Room 404 in Building-1 with evacuees",
        "texts": [
            " Contributed by Computers and Information Division of ASME for publication in the JOURNAL OF COMPUTING AND INFORMATION SCIENCE IN ENGINEERING. Manuscript received November 5, 2011; final manuscript received March 1, 2012; published online June 7, 2012. Assoc. Editor: Bahram Ravani. Journal of Computing and Information Science in Engineering SEPTEMBER 2012, Vol. 12 / 031001-1 Copyright VC 2012 by ASME Downloaded From: http://computingengineering.asmedigitalcollection.asme.org/ on 08/22/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 4 shows room 404 on the fourth floor of building 1 along with the related evacuee. The room model involves structural parameters, such as length and width, and contains the personal information of the evacuee inside the room. For example, a total of 27 evacuee and their names are registered in room 404. In the case of a mass-evacuation, a number of evacuation data must be processed simultaneously. Hence, for computational efficiency, only the data that are necessary for evacuating individuals are processed at each of the current location",
            " Each facility model has its local coordinate frame at its origin, which is related to the global coordinate frame by Rg. The local coordinate frames are expressed by a homogeneous transformation matrix gTl [ R4 4. Before evacuation, students line up in the designated location in the classroom. The students then rush together into the narrow exit when the evacuation is announced. For this reason, as evacuation goals, the local coordinate frames Rexit_left and Rexit_right are located at the sides of the room exit (Fig. 4). 2.2 Evacuation Model for Simulation 2.2.1 Algorithm of the Evacuation Model. Figure 5 shows the occupation-inspection space moving with an evacuee. The space consists of three adjacent blocks, each with depth B, width W, and height H (not shown). This space is generated on an ad hoc basis accompanying the evacuee motion. Student A, for example, takes a step forward if no evacuee is located in front of student A, or looks to the right if another evacuee is in front of student A. Simultaneously, student A looks to the left to find a vacant space into which to step forward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure14.5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure14.5-1.png",
        "caption": "Fig. 14.5 A free body diagram of the aerodynamic forces on a maneuvering missile. The diagram includes all terms necessary for analysis of the free response of the short-period approximation",
        "texts": [
            " Though this accurately describes the kinematics of the engagement, it neglects the dynamics of the missile and creates an impractical control vector. As was shown in the engagement kinematics, a missile must accelerate (aMn ) to intercept a maneuvering target. For an endoatmospheric missile, body rotation is the most efficient means of generating the necessary forces to accelerate the vehicle and counteract the target\u2019s maneuvering. Therefore, the rotational missile dynamics must be modeled accurately. As was the case for the engagement kinematics, motion is constrained to the longitudinal plane. Consider the rotated missile body in Fig. 14.5, where x, z is the coordinate system attached to the body of the vehicle, U is the primary axis attached to the wind frame, and X is the primary axis attached to the same inertial frame of reference used in the engagement kinematics. The angle of attack \u03b1, flight path angle \u03b3 , and pitch angle \u03b8 , track the orientation of these coordinate systems relative to each other. In this scenario, the angle of attack \u03b1 leads to an aerodynamic force F on the body at the vehicle\u2019s center of pressure. This force is then represented as a set of forces (Fx and Fz) and a moment M at the center of mass",
            " The free response of the system, that is, x\u0307 = Ax such that u = 0, is dependent on the stability derivative terms. Stability and frequency of the short period is dictated by the value of Cm\u03b1 . A negative value indicates a stable mode, and is confirmed by the center of gravity\u2019s forward position with respect to the vehicle\u2019s center or pressure The mode\u2019s damping is reliant on the magnitude of Cmq , and the vehicle\u2019s acceleration effectiveness is determined by Cz\u03b1 . All of the terms required for free response analysis are conceptualized in Fig. 14.5. The forced response of the system (u = 0) requires the stability derivative terms from Fig. 14.5 and the control derivatives from Fig. 14.6. It is important to think of the fins as a moment generators and not force generators. The control derivative term Cm\u03b4 is responsible for producing that moment. It is a product of the force generated at the fin and the moment arm from fin center of pressure to the vehicle\u2019s center of gravity. Provided there is a fixed moment arm length, increasing the surface areas of the actuation fins will lead to greater rotational authority. However, this comes at a cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002912_s12555-020-0049-x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002912_s12555-020-0049-x-Figure1-1.png",
        "caption": "Fig. 1. ARP CAD model.",
        "texts": [
            " The remainder of this manuscript is organized as follows: Section 2 presents the dynamic model of the sit\u2013 stand mechanism. In Section 3, the design of controllers is explicated. The simulation and its results are discussed in Section 4, and Section 5 summarizes the conclusions. 2. DYNAMIC MODELLING OF SIT-STAND MECHANISM A dynamic model is necessary for the simulation and analysis of a system, because it is the first step toward control implementation. The mathematical model of the STS mechanism describes its motion between the sitting and standing positions. The ARP CAD presented in Fig. 1 includes two parts: an STS mechanism to support the shift of a patient from a sitting to a standing position and vice versa and the mobile platform controlled by a joystick. Khan and Kausar discussed the kinematic modelling of ARP [30]. The robot mechanism is considered to be the revolute\u2013prismatic\u2013revolute (RPR) mechanism shown in Fig. 2. The Lagrange approach implements the equation of motion of the system [31]. This dynamic equation can be expressed as (1): M(\u03b8)\u03b8\u0308 +C(\u03b8 , \u03b8\u0307)\u03b8\u0307 +N(\u03b8 , \u03b8\u0307) = \u03c4, (1) where \u03c4 , M, C and N are actuator torque vector, mass matrix, Coriolis matrix and gravity term respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001859_j.promfg.2020.10.106-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001859_j.promfg.2020.10.106-Figure4-1.png",
        "caption": "Fig. 4. Sieve: (a) Assembly model, (b) Assembly photo, (c) Cam, (d) Connecting rod.",
        "texts": [
            " Furthermore, by adjusting the speed at which the motor rotates powder inflow control is affected (powder weight per unit time). These two controls combined optimize the time needed to refill the sieve between layers. This step is necessary in order to better exploit raw material because used powder\u2019s properties deteriorate every time it is used due to the temperature of the heated table. This is also time-efficient, since as little time as possible is spent in the transition between building phases of the printing. Fig. 3. Powder doser drum components. The complete assembly is shown in Fig. 4(a)-(b). Palindromic motion of a mesh is relied on in order to sieve the powder raw material. The mesh is interchangeable, which enables the selection of various mesh sizes to achieve the desired size distribution of powder particles. This capability is of paramount importance, because a different size distribution of powder grains causes the rheological characteristics of the powder to change thus affecting the packing density and surface roughness of the spread powder layer. The palindromic motion of the sieve is provided by a stepper motor which defines a cam, see Fig. 4(c), rotating in the groove of a connecting rod, see Fig. 4(d), thereby converting the rotary into an oscillatory linear motion. The sieve\u2019s rod is sliding inside bushings to minimize friction. While sieving, the granules that surpass the size limit as well as larger agglomerates of granules are withheld in the sieving tank. If these were deposited on the powder bed, they would be subsequently dragged along the powder layer by the doctor blade or roller, thereby creating deep grooves on the powder layer and ultimately rendering it useless. 758 Panagiotis Avrampos et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001884_1350650120975519-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001884_1350650120975519-Figure1-1.png",
        "caption": "Figure 1. Geometric structure of spherical plain bearings. (a) Whole structure (b) Cross section.",
        "texts": [
            " In this paper, the distribution of contact pressure between the inner and the outer rings of bearings under normal loads is investigated with the help of finite element methods. A corresponding analysis is performed based on the conformal contact theory and the elastic foundation model to reveal the effects of tribology on the contact pressure distribution. At last, the axial contact pressure is specially discussed to clarify the underlying mechanism of the free-edge effects. The geometric structure of spherical plain bearings employed in this study is shown in Figure 1 with the corresponding parameters listed in Table 1, where d, D, B, C, dk and Ca represent the inner diameter, outer diameter, inner width, spherical diameter and related dynamic load of the bearings, respectively. The contact surfaces consist of the inner spherical surface of the outer ring and the outer of the inner ring, as depicted by the red curves in Figure 1(b). It is noteworthy that the interval DR of contact surfaces has a significant effect on the distribution of contact pressure.18 When the interval is considerably small, the contact pressure concentrates near the edges, moving towards the center with the increase of DR. In this study, the interval is set to be zero to study the stress concentration on the free boundary, which means that the radius of both contact surfaces are equal. According to the conformal contact theory,15 the circumferential contact between the inner and the outer rings under normal load is an axis-symmetric elastic issue as shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000274_icems.2019.8921504-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000274_icems.2019.8921504-Figure4-1.png",
        "caption": "Fig. 4. Flux distribution of four topologies. (a) Model 1. (b) Model 2. (c) Model 3. (d) Model 4.",
        "texts": [
            "3(b) are same to analysis results in Table II. This work was supported in part by National Key R&D Program of China under Project 2017YFB0203603, in part by National Natural Science Foundation of China under Project 51637003, 51607046, and 51777013. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE Usually the (Qs-Qpm) is small (2 in this paper), so its pole pitch and flux per pole are high. This component is needed in PMVM and the pole pairs of armature field is (Qs-Qpm). III. COMPARISON OF TOPOLOGIES WITH INITIAL PARAMETERS Fig.4 shows the flux distribution of four topologies. Due to interior PMs, leakage flux in Model 3 and Model 4 are bigger than the other two topologies although magnetic barriers are used, which could also be proved by the amplitude of (Qs-Qpm) PM pairs component in Fig.3. With initial parameters, Model 1 has the highest amplitude of needed component. Amplitude of Model 2 is lower than Model 1 because the proportion of tangentially magnetized PM is high (50%). The no-load back EMF of four topologies are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001656_etfa46521.2020.9211891-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001656_etfa46521.2020.9211891-Figure5-1.png",
        "caption": "Fig. 5: Gripper with force/torque sensor and connector",
        "texts": [
            " Localization of the connector is supposed to be achieved by means of a controlled motion using contact force feedback. The automated testing system will be equipped with a 6 DOF industrial robot. Fig. 4 shows the experimental setup with ACF and two Comau Racer3. Notice that only one of the robots is used. The robot is equipped with a JR3 67M force/torque sensor mounted at its end effector. This sensor provides the necessary force feedback. The designed gripper for the plugging procedures with end effector frame FE is shown in Fig. 5. The proposed approach comprises the following steps: 1) Force controlled approaching towards the plug so to get the plug in contact with the mating face. 2) Coordinated translational motion of plug while remain- ing in contact with the mating face so to match the cylindrical surfaces of the plug with that of the jack. 1466 Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 02,2020 at 11:44:48 UTC from IEEE Xplore. Restrictions apply. 3) Partial insertion of the plug until the inverse-polarity protection is contacted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000686_mepcon47431.2019.9008048-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000686_mepcon47431.2019.9008048-Figure1-1.png",
        "caption": "Fig. 1. Solid model of one-third of the proposed machine.",
        "texts": [
            " For faster FE analysis, one-third of the machine is modeled, which corresponds to 12 stator slots and 15 rotor bars. In order to take into account the remaining two-third part of the machine geometry, even periodic boundary condition is employed with a repetition angle of 120\u00b0. Even periodic boundary condition is selected based on the symmetry of flux paths in machine geometry, since the flux pattern in core doesn\u2019t change from one-third repeating section to the next. For periodic boundary conditions, it is important to ensure that phase coil voltage should also be reduced to one-third of phase voltage. Fig.1, shows solid model of one-third of the considered motor drawn in 2-D plane. Basic machine dimensions and parameters are summarized in Table I. Initially, the model is energized at full load conditions and flux distribution is shown in Fig.2. Rotor bar currents are taken from full load analysis to be used later for conducting RST. It is important to emphasize that, the following parameter estimation procedure can readily be applied to an induction machine with any number of phases greater than 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001805_ecce44975.2020.9236039-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001805_ecce44975.2020.9236039-Figure1-1.png",
        "caption": "Fig. 1. Examples of general forms of SRMs.",
        "texts": [
            " The new contribution of the proposed method, the triple zero is achieved with the same efficiency as the general driving method of SRM by utilizing the control flexibility effectively. This paper is organized as follows; first the configuration of the five-phase SRM and its control flexibilities are explained. Next, the derivation method of the ideal current waveform to achieve the zero torque ripple, zero DC current ripple, and zero radial force ripple simultaneously is explained. Finally, the finite-element analysis and experimental results are shown to validate the proposed method. II. TRIPLE ZERO WITH FIVE-PHASE SRM Fig.1 depicts the motor diagrams, whereas Table.1 shows the motor parameters of the three-phase SRM and the fiveThis work is financially supported by JSPS KAKENHI Grant Number 19K21968. 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 4703 Authorized licensed use limited to: University of Gothenburg. Downloaded on December 18,2020 at 11:37:10 UTC from IEEE Xplore. Restrictions apply. phase SRM. For comparative verification, the five-phase SRM is designed to have the same specification and same size as our three-phase 18S/12P SRM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000703_s12206-020-0122-7-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000703_s12206-020-0122-7-Figure17-1.png",
        "caption": "Fig. 17. Analysis result for deformation.",
        "texts": [
            " 12; Bolt-fastening component 1, fastening component 1 - fastening component 2, fastening component 2 - fastening component 3, fastening component 3 - nut, bolt - nut. The coefficient of friction was obtained by the data derived from the test. Tightening torque of the same magnitude as the test is applied to the bolt head. In addition, the nut is constrained by all the degrees of freedom except for the axial length of the bolt. At this time, the fastening force is derived through the contact force among the fastening components. Table 3 shows the value of clamping force through the experiment and finite element analysis. Fig. 17 shows the displacement of the bonded body after the fastening analysis. The displacement of the clamping surface is maximum 0.033 mm and the contact force (clamping force) is 51.9 kN. The purpose of the analysis is to increase the reliability of the analysis results by comparing and verifying between the analysis and the test. Then, it is possible to deduce the accurate loosening load at the time when the unfastening phenomenon occurs, and to save time and cost. In addition, the analysis is intended to facilitate continuous expansion of future research by deriving various data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001528_0954407020951318-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001528_0954407020951318-Figure4-1.png",
        "caption": "Figure 4. Arc-spring friction diagram.",
        "texts": [
            " In summary, a nonlinear elastic model of the twostage piecewise stiffness of the DMF can be obtained as follows: K(u)= ka2(u u1)+ ka1(u1 u0) u . u1 ka1(u u0) u0 \\ u \\ u1 0 u0 \\ u \\ u0 ka1(u+ u0) u1 \\ u \\ u0 ka2(u+ u1) ka1(u1 u0) u \\ u1 8>>< >>: \u00f01\u00de where u is the relative twist angle of the first mass and the second mass,K(u) is nonlinear stiffness function.The two-stage piecewise torsional stiffness characteristic is shown in Figure 3. Friction damping modeling of arc-spring The working process of the arc-spring and the sheath generate friction damping torque. The friction damping torque is composed of two parts, as shown in Figure 4. The first part of the friction force is generated by the radial component of the arc spring force. The other part of the friction is generated by the centrifugal force during the rotation of the DMF. Friction damping torque generated by the radial component. In this paper, the discrete element method is used to analyze the force acted on the arc-spring. The total mass of the arc-spring is m, which is divided into n units. The mass of each unit is mi (i = 1, 2, 3...n), and the unit stiffness is ki"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001368_speedam48782.2020.9161859-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001368_speedam48782.2020.9161859-Figure1-1.png",
        "caption": "Fig. 1. Top view of induction motor",
        "texts": [
            "00 \u00a92020 IEEE 157 Authorized licensed use limited to: Cornell University Library. Downloaded on August 29,2020 at 16:24:07 UTC from IEEE Xplore. Restrictions apply. III. EXPERIMENTAL APPARATUS A. Induction motor To test the motor, one microphone was placed in front of the motor, in line with its shaft, and three accelerometers are placed in significant positions: 1) over the supporting base, 2) in line with the shaft and 3) at the frame back. The top view of the acoustic test rig is reported in Fig. 1. The side view of the test rig, showing the same sensors, is reported in Fig. 2: The characteristic parameters of the cage induction motor used for experiments are given in Tab.III-A: Specification Value Unit P 0,75 kW cos\u03d5 0,8 V\u0394 230 V Vy 380 V I\u0394 3,25 A Iy 1,88 A n 1395 rpm The electric drive is controlled with a Space-Vector Pulse-Width Modulation (SVPWM) technique, implemented in a field programmable gate array (FPGA) embedded in a control system NI-PXI built by National Instruments\u00ae. The control scheme is implemented in the Labview\u00aeFPGA environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001861_1350650120972499-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001861_1350650120972499-Figure3-1.png",
        "caption": "Figure 3. Schematic of the slipper\u2019s macro trajectory on the swash plate.",
        "texts": [
            " Before deriving the Reynolds equation of a slipper bearing in a cylindrical coordinate system. The lubricating oil film is regarded as Newtoian, imcompressible and hydrodynamic flow without solid surface deformation effect. The fluid film pressure is governed by the Reynolds equation for the slipper bearing and is expressed as18 1 r @ @r qrh3 l @p @r ! \u00fe 1 r2 @ @h qh3 l @p @h ! \u00bc 6vTr @qh @r \u00fe 6 vTh r \u00fe xss @qh @h \u00fe 12 @qh @t (4) When the slipper slides on the swash plate, the transitional velocity of slipper includes radial velocity and tangential velocity. Figure 3 shows the schematic of the slipper\u2019s macro trajectory on the swash plate. The radial and tangential velocities of slipper are respectively written as19 vTr \u00bc xRs \u00bc cosb ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe tan2cos/ p cos2/\u00fe cos2bsin2/ xRpsinh (5) vTh \u00bc xRs \u00bc cosb ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe tan2cos/ p cos2/\u00fe cos2bsin2/ xRpcosh (6) Figure 4 shows the boundary conditions of a textured slipper. The boundary conditions of the textured slipper bearing can be expressed as: p h; r \u00bc r0\u00f0 \u00de \u00bc ps; p h; r \u00bc R\u00f0 \u00de \u00bc 101325 Pa (7) The geometric characteristics of the sealing belt of the slipper affect the pressure periodically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001120_j.jfranklin.2020.05.018-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001120_j.jfranklin.2020.05.018-Figure1-1.png",
        "caption": "Fig. 1. A scene of topology construction.",
        "texts": [
            " In contrast, the inner attitude loop can be used to track the desired angles. The tracking problem of inner attitude loop is well established in [18 , 19] , so we mainly focus on designing the position loop in this paper. Considering Eq. (8) , the reduced equations of position loop can be written as { \u02d9 pi = v i \u02d9 vi = u p i + d F , (9) w u 3 l t a q t [ t o S D c D s here u p i \u2208 R 3 is virtual control inputs for the position dynamics of the i agent. Then, these inputs can be written as p i = 1 m U 1 i R e 3 \u2212 g e 3 . (10) . Problem Description Fig. 1 shows a classical location of hunting that abstracts from the dolphin hunting the antern fish. Consider a target whose state is (p t , v t , a t ) . The control objective in this paper is o design a distributed control protocol for the MQS such that the quadrotors can constitute n autonomous formation surrounding the target and gradually captures it. In this process, the uadrotors only receive the relative distance vector q k ( t ) and bearing information c k ( t ), and hey spirally shrink the radius with rotary velocity \u03c9 k ( t ) to prevent the target from escaping",
            " Meanwhile, we also found that the larger the upper bound is, we could get smaller eigenvalues of A \u2212 KC . However, a feasible gain K may not exist when the range of time delay is too large. Therefore, the observer gain matrix K needs to be verified by simulation results. 6. Simulations In order to verify the effectiveness of the hunting protocol and delay state observer, MATLAB-R2017a is used to build the model and run the simulation. The initial parameters of simulation are set as shown in Table 1 , and all the models are the same. \u2223\u2223D F xy \u2223\u2223 \u2264 0. 1 and \u2223\u2223D F z \u2223\u2223 \u2264 0. 2. According to Fig. 1 , initial expected bearing constraints between agents are p \u2032 12 = \u2212p \u2032 21 = [1 , 0, 0] T p \u2032 23 = \u2212p \u2032 32 = [ 1 2 , \u2212 \u221a 3 2 , 0] T , p \u2032 34 = \u2212p \u2032 43 = [ \u22121 2 , \u2212 \u221a 3 2 , 0] T , p \u2032 45 = \u2212p \u2032 54 = [ \u22121 , 0, 0] T , p \u2032 56 = \u2212p \u2032 65 = [ 1 2 , \u2212 \u221a 3 2 , 0] T , p \u2032 61 = \u2212p \u2032 16 = [ 1 2 , \u221a 3 2 , 0] T . (53) P P Based on Eq. (3) , the definition of projection operator, we can obtain P \u2032 12 (0) = P \u2032 45 (0) = P \u2032 3 t (0) = P \u2032 6 t (0) = \u23a1 \u23a2 \u23a3 0 0 0 0 1 0 0 0 1 \u23a4 \u23a5 \u23a6 , \u2032 23 (0) = P \u2032 56 (0) = P \u2032 1 t (0) = P \u2032 4t (0) = \u23a1 \u23a2 \u23a2 \u23a3 3 4 \u221a 3 4 0 \u221a 3 4 3 4 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 , \u2032 34 (0) = P \u2032 61 (0) = P \u2032 2t (0) = P \u2032 5 t (0) = \u23a1 \u23a2 \u23a2 \u23a3 3 4 \u2212 \u221a 3 4 0 \u2212 \u221a 3 4 3 4 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 , P \u2032 i j (t ) = R(\u03c9t ) P \u2032 i j (0) R (\u03c9t ) T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002935_iemdc47953.2021.9449597-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002935_iemdc47953.2021.9449597-Figure2-1.png",
        "caption": "Fig. 2. E-core radial AMB. A flux density contour plot for the maximum control current on the y-axis electromagnets and the bias current present.",
        "texts": [
            " Restrictions apply. When comparing the twin bearingless motor with the traditional AMB rotor, the axial bearing design remains the same. The application process calculation gives the boundary conditions for the design of low- and high-pressure compressor wheels located at opposite rotor ends. The force imbalance resulting during the operation together with gravity define the radial and axial force loads in vertical rotor. To carry radial loads of the rotor an E-core radial AMB was selected for SPM-AMB rotors. Fig. 2 shows the E-core radial AMB and its flux density contour plot. A non-saturated design provides a relatively linear force at the expense of maximum force capacity. The E-core structure decreases the stator and rotor yoke thicknesses compared with the C-core structure. The main radial AMB parameters are shown in Table III. The SPM motor is designed as 160 kW at 30000 r/min. However, the operational speed range starts from 17000 r/min. To minimize the losses, a 2-pole structure and a distributed winding are applied"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure8-1.png",
        "caption": "Fig. 8. Composition diagram of the right device",
        "texts": [
            " Here, we will introduce in detail the design of the device on the side affected by the angle steel structure of the tower, and will not explain the device on the side that is not affected by the angle steel. The right side device and the left side device have the same appearance outline size, and are composed of an top surface, a bottom surface, a front side surface, and two side surfaces. Among them, the bottom, two sides, and the front side are affected by the angle steel structure. There will be grid gaps in the contact part of the angle steel, as shown in Fig.8, so that the device can bypass the angle steel and fit the left device, thus avoiding a gap between the left and right devices after splicing. It can be seen from the Fig.9. that the front side of the bird prevention device on the right side of the middle phase is affected by the V-shaped angle steel. The entire front side is divided into three parts, of which the two triangular mesh surfaces are symmetrical. In the middle is a symmetrical hexagonal mesh surface. 978-1-7281-6246-1/20/$31.00 \u00a92020 IEEE 810 Authorized licensed use limited to: University of Canberra"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001916_icem49940.2020.9270720-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001916_icem49940.2020.9270720-Figure7-1.png",
        "caption": "Fig. 7: 2D FEM model of motor Linix in Ansys-Maxwell",
        "texts": [
            " 1241 Authorized licensed use limited to: UNIVERSITY OF NEW MEXICO. Downloaded on May 15,2021 at 03:47:51 UTC from IEEE Xplore. Restrictions apply. Fig. 6 shows construction of 3-phase PMSM Linix 45ZWN24-40 and tab.1 contains its parameters. We tried to create the most accurate and the most reliable 2D model of this motor in Finite Element Method simulation software ANSYS-Maxwell. Then the FEM 2D model of the motor was coupled with simulation of Field Oriented Control in ANSYS-Simplorer(Twin Builder) as is shown in fig. 8 to simulate transient state. Fig. 7 shows 2D FEM model of used motor Linix 45ZWN24-40 in ANSYS-Maxwell software. Simulation model in fig. 8 composes of speed loop with PI controller tuned for 15 Hz bandwidth. Output from speed PI controller is required q-axis current. Required d-axis current is set for 0 Amps. Then required dq currents are transformed to abc reference frame. Conventional dq current loops in standard FOC are in our simulation replaced by three current sources controlled by required abc currents to eliminate impact of limited current loops bandwidth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure11-1.png",
        "caption": "Figure 11. Moment 3: (a) proposed model (b) FE model.",
        "texts": [],
        "surrounding_texts": [
            "Figure 8. Finite element model in ABAQUS. will lead to a decrease in computational speed, so this article selects the number of pieces as 20. Based on the above results, the mean error of contact force along axis is 7.1% solved by the sum of the error on each slice divided by the number of slices, and the mean error of contact force of one tooth is 4.2% solved by the sum of the error at each moment divided by the number of moment sampling points, as shown in Figures 13 and 14. The comparison results verified the proposed model and the result of the proposed model well reveals the pattern of contact force distribution of beveloid gear pair. Table 3. Status of each moment. Moment Rotation angle Mesh status Moment 1 28 Two teeth in contact Moment 2 68 One tooth in contact Moment 3 108 Two teeth in contact Figure 12. Effect of the number of pieces."
        ]
    },
    {
        "image_filename": "designv11_63_0003074_s11804-021-00201-6-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003074_s11804-021-00201-6-Figure1-1.png",
        "caption": "Figure 1 Sketch of air cushion vehicles model",
        "texts": [
            " Compared with the other studies that rely on the measurability of states of the ACV, a sensorless approach is derived in this paper. The neural observer is designed to identify the nonlinearity with high accuracy and guarantee the convergence of the observer to zero. In addition, the terminal sliding-mode controller based on the proposed observer is developed to satisfy the closed-loop system stability. The nonlinear dynamics adopted in this paper has been derived and identified in Liang et al. (2019). The equations of velocity components as shown in Figure 1 are presented below: u\u0307 \u00bc \u2212m\u22121du0 sgn u\u2212m\u22121duu\u00fe m\u22121bTTcos\u03b8\u00fe vr v \u00bc \u2212m\u22121dv0 sgn v\u2212m\u22121dvv\u00fe m\u22121bTTsin\u03b8\u2212ur r\u0307 \u00bc \u2212J\u22121dr0 sgn r\u2212J\u22121drr\u2212J\u22121abTTsin\u03b8 8< : \u00f01\u00de where u and v are the surge and sway velocities; r is the angular velocity; {du0, du, dv0, dv, dr0, dr}\u2208R are the friction and drag coefficients; a is the length of the arm from the center of mass to the rudder surface; T is the thrust force; and \u03b8 is the rudder angle; the coefficient bT scales the thrust input from [0, 1] to force in Newton. The model is actuated by the thrust force T and rudder angle \u03b8, which are generated by propellers subjected to velocity drag forces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000004_aim.2019.8868575-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000004_aim.2019.8868575-Figure9-1.png",
        "caption": "Fig. 9. Presliced topology optimized geometry and conductive traces, with dual extruders assigned and tree-like supports (left); coupled 3D printed conductive and non-conductive housing (right)",
        "texts": [
            " This result creates the premise for future optimization which should exploit the percolation effect [16] and electrical anisotropy of the conductive paths, thus optimizing the current flow through minimizing the resistance within the printed conductive traces. The obtained topology optimized geometry is then sliced using FlashPrint, the dedicated slicer of the Flashforge Dreamer 3D printers. The conductive traces are printed simultaneously in contact with the non-conductive housing, as depicted in Figure 9 along with the physical, 3D printed device. The assembly process consists of inserting the motor and the battery. The selected Tamyia RC300-FT-14270 DC motor represents a typical off-the-shelf component. It exhibits the application convenient low-current characteristics [23] of 25 mA at 1.5 V, as well as solid terminals which can be contacted against the conductive traces. The battery is a standard 9V battery from Q-Batteries [20] and is inserted into the battery holder, as can be seen in Figure 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure16-1.png",
        "caption": "Fig. 16. Inverted umbrella-shaped foldable anti-bird thorn",
        "texts": [],
        "surrounding_texts": [
            "The dumpable structure is to pour the bird thorn to one side after the anti-bird thorn is folded, which provides more work space for maintenance. The anti-bird spur and the base part are connected by screws. When the anti-bird device needs to be tipped over, the screw of the connection part can be loosened with a screw wrench to tilt the bird spur to one side to avoid affecting operation and maintenance. After the overhaul is over, the bird spines will be straightened and fixed, and then the bird spines can be opened."
        ]
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure3-1.png",
        "caption": "Fig. 3. Views of windows manifold (a) connection to commutator, (b) cross-section in the middle, (c) connection to rotor.",
        "texts": [
            " 2, enters from one of the ports in ports block, 1, and it passes through the slots before being supplied to the outside of the commutator, 6. The outlet flow, represented in blue, is provided to the inside of the commutator by the windows manifold, 5. The inside of the commutator is directly connected to the through hole of the spline joint, 3, through a slot in the latter. The through hole of the spline joint is connected to the other port of the ports block. Therefore, the commutator separates the inlet and outlet environments. Three views of the windows manifold are presented in Fig. 3. In these views, the diametrically outer slots represent the passages through which the studs pass through. The flow enters or leaves through the inner slots presented in Fig. 3(a), as dictated by the commutator. Fig. 3(c) represents the view of windowsmanifold when viewed from the rotor side. Each one of the inner slots represented in Fig. 3(c) is connected to each of the TSV in the rotor-roller set. The commutator and rotor are connected by the floating spline joint (Fig. 2) and thus they move synchronously. For the continuous operation of the orbit motor, high pressure fluid is provided at a different angle to the TSV, as represented by the cross-sectional view of Fig. 3(b). Fig. 4 represents the windows under consideration along with the commutator (colored in gray) over an orbit rotation. The commutator performs the orbiting action, subjecting each of the windows in the windows manifold, 5 (Fig. 2), to the inside or outside of itself. From Fig. 4, it can be observed that the windows are exposed to the inside of the commutator or the outside which correspond to the outlet (colored in blue) and inlet environments (colored in red) respectively. For a certain duration of angle, the window area is completely enveloped by the commutator thus providing zero cross-porting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001125_j.matpr.2020.05.084-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001125_j.matpr.2020.05.084-Figure2-1.png",
        "caption": "Fig. 2. Finite element mesh of 0.5 mm size and tetrahedron shape for horizontal S-shaped model.",
        "texts": [
            " The mechanical behaviour of all the three structural forms (circular, horizontal and elliptical S-shape) were simulated for compressive loading using the ANSYS Workbench 2019. The structures were fixed from one end (+X direction), and a displacement based compressive load of 0.1 mm was applied from the other end (-X direction) leaving Y and Z directions free. The structural deformations were simulated by finite element analysis for linear material conditions, choosing stainless steel 316L as the base material. The uniform FEA mesh of the tetrahedron shape and 0.5 mm of size was generated for all the models. Fig. 2 is depicting the applied boundary conditions, and image in the inset shows the tetrahedral discretisation more closely for horizontal elliptical form. The size of the finite element mesh at around 0.5 mm was obtained based on a mesh convergence analysis result indicating the lateral displacement to converge on a fixed value when the element size was iteratively reduced to around 0.50 mm. Fig. 3 shows the stress distribution patterns and auxetic responses emanating in the three different S-shaped model variants analysed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001023_iccre49379.2020.9096444-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001023_iccre49379.2020.9096444-Figure1-1.png",
        "caption": "Figure 1 Arrangement of the 3 DOF",
        "texts": [],
        "surrounding_texts": [
            "training is one of the best ways of\ntraining badminton. But it is difficult to practice badminton without the assistance of another skilled player and coaches cannot train large number of players during multi-shuttle training session. To avoid these problems multi-shuttle training machines have been introduced. But these machines are expensive and can be further improved to gain more accuracy and to reduce shuttle damage. This paper describes a badminton shuttlecock feeding machine which was designed and developed in order to reproduce the actual shots. Based on the previous researches shuttlecock trajectory was simulated for four basic badminton shots and obtained the initial ejecting parameters. Then a testing apparatus was developed to test and select suitable ejecting mechanism, to test error and test damage\nduring the ejection. According to the testing results, optimum ejecting mechanism was selected and further the torque calculation was carried out for the dispenser and tilting of roller section. In this paper, a description is given on the prototype designing and development, along with the development of a controlling architecture to control the shot parameters. Finally the fabricated\nprototype was tested and validated with the use of simulation results and experiments were carried out to find out\nthe degree of damage in the shuttlecocks. In the near future the fabricated machine will be refined with the use of experimental results.\nKeywords\u2014Badminton; multi-shuttle training; shuttlecock\nfeeding; shuttlecock trajectory; shuttlecock ejecting machines;\nI.\nINTRODUCTION\nBadminton shuttlecock feeding machine which is always referred to multi-shuttle training machine, is a very important research area since its complexity is higher than other similar machines. The use of these machines are very important for both coaches and players since it is difficult to practice badminton without assistance of another skilled player and coaches cannot train large number of players during multi-shuttle training session. To avoid these problems multi-shuttle training machines have been introduced. But as for the current status of these machines, they can be further improved to gain more accuracy and to reduce shuttle damage.\nSimilar to the concept of multi-shuttle training machine, the concept of multi-ball machine is very popular among the sports such as cricket, tennis, baseball, volleyball and etc. Design and controllability of these machines are not that much difficult compared to multishuttle training machines since the shape of the object is not complex as the shuttlecock. Over past 30 years many researches have been carried out to determine the\nparameters that affects the trajectory of the shuttlecocks and to find optimum conditions to achieve best of the game. According to previous researches shuttlecocks generate high drag on the feathers or rubber skirts of the shuttlecock due to high initial speeds [1].\nWhen shuttlecock\nis flying it will mainly subject two forces, which are drag force and gravity force. Drag force depends on the Reynolds\nnumber and shuttle rotating speed increased with increasing Reynolds number [2].\nBased on these factors Chen, Pan and Y.J. Chen 2009 have developed an\nequation for the trajectory motion of the shuttlecock [3].\nHowever the type of shuttlecock, whether the fact that it\u2019s a synthetic one or a goose feather one greatly affects the trajectory path [1].\nOver the course of the past 100 years several inventors have introduced different types of mechanisms to successfully achieve the high speed shuttlecock ejecting. One of the early method was to use compressed air for ejecting shuttlecocks\n[4]. The mechanism involved complex loading mechanism and it didn\u2019t give\nfull control on shuttle ejecting speed. Another early design was to use an impact method to eject shuttlecocks\n[5]. This design also involved complex mechanisms and also its capability to produce higher speeds was low. Striking mechanism which was used here had a high degree of possibility to damage the shuttlecock. Another design which occupied a swing arm mechanism for ejecting shuttles was introduced by Otmar Schall in 1989 [6].\nThis design also had the same problems such as lower ejecting speed and mechanical complexity. Apart from above designs another mechanism which involved spring load mechanism was introduced and tested by the researchers\nYusif and Kok in 2011 [7]. The limiting factor in this mechanism is when it comes to the production of high speeds since it uses\na spring. Among all the introduced mechanisms most popular mechanism is to use two rollers for ejectors. This was the same mechanism which used in ball\nfeeding machines in other related sports. First design which occupied the two roller mechanism was introduced by Jonathan Taryoto in 2004 [8]. This two roller mechanism had greater advantages compared to other mechanisms.\nCurrently number of commercial machines have been introduced which utilize the same two roller mechanism. Knight trainer is one of the earliest commercially successful badminton multi-shuttle training machines\n[9]. Although it had limited programming capabilities it was capable of training the players to a certain extend. The newest version of this trainer which was named after Knight Trainer Pro, is more advanced training machine\n73978-1-7281-6791-6/20/$31.00 \u00a92020 IEEE\nAuthorized licensed use limited to: Murdoch University. Downloaded on June 13,2020 at 23:52:11 UTC from IEEE Xplore. Restrictions apply.",
            "with very good programming capabilities [10]. This machine has ability\nto produce different shots at a sequence. It is also\ncapable of feeding different shots with user defined parameters. Apollo trainer is yet another advanced multi-shuttle training machine which has been commercially introduced\n[11]. It has a high power of programming capabilities in order to utilize placement sequences and speeds. Apart from those machines several other products have been introduced to the market with lower functionalities\n[12, 13]. One of the major disadvantages with these commercial machines would be their high\nprice. Additionally, the training method used in some of these products, is developed primarily focusing on locating the shuttlecock at different points. It would be quite different from what is experienced with an opponent.\nThis mechanism still has number of research areas to be carried out depending on the roller angles, gripping position and other parameters. The research paper which was published as \u201cA study on Projection Performance of Roller type Badminton machine and its Optimization\u201d has enclosed several effective parameters to be used in this mechanism\n[14]. The researchers have simulated the stress distributions on the rollers and the shuttlecock and obtained the most effective values for the parameters in the rollers. Apart from that they have tested the accuracy of the mechanism with a developed prototype. The results of this research was concluded with optimum roller parameters. However they have used a different shuttlecock gripping positions for ejecting and the rollers are very large. But their results are adequate and promising.\nIn addition to these mechanisms, a novel highspeed and lightweight Humanoid Robot Arm\nhas been introduced by a Japanese research group [15]. This robot arm is operated using pneumatic actuators and it is designed to operate as an intelligent opponent.\nAs mentioned earlier shuttlecock has unique properties compared to balls in other sports. Considering these properties and indoor air properties, trajectory equations have been introduced. Using these equations it is possible to\nsimulate the path of the shuttlecocks for different initial conditions.\nThe relationship between Reynolds number with speed of flight and effect of buoyancy force during the shuttlecock flight has also\nbeen experimentally analyzed by previous researchers [16].\nAs for the badminton sport,\nthere are six basic shots and they are drop, net-lift, smash, drive, defensive clear and attack clear [3]. In this research the target was\nto reproduce four shots clear, net lift, smash and drive, in the prototype. Using the trajectory equations and simulation results,\nsuitable initial parameters can be obtained.\nThe section II will introduce the system parameter identification for to process in the design. In section III, the modified trajectory equations have been simulated\nand obtained the best suitable initial parameter values for given four shots. Then, section IV discusses\nabout an experiment carried out in search of\na suitable ejecting mechanism by developing a testing apparatus. In section V\na short description is presented on the developed\nmodel of the prototype with the suitable ejecting mechanism obtained from section IV. In section VI, implemented\ncontrolling architecture is described. The progress and the current state of the development of the prototype is presented in the section VII. The section VIII discusses about the\nexperiments and results which were carried out to evaluate the design. Finally the section IX presents\nsome conclusions based on the experimental results and future directions of the project.\nII. SYSTEM PARAMETER IDENTIFICATION\nWhen it comes to analysis of badminton shots created by an actual badminton player, it can be observed that the parameters involved when the contact happens between the shuttlecock and the badminton racket are\nimportant. These parameters can be classified in terms of position wise, direction wise and contact speed wise. During the project it was identified that the following variable parameters will only be sufficient enough to reproduce the badminton shots.\nTranslation along Y-axis \u2013 This will allow to eject shuttlecocks at different heights.\nTilting of the shuttlecock ejecting unit \u2013 This will allow to eject shuttlecocks at different angles to produce shots, such as net-lift, drive, smash.\nPanning of the shuttlecock ejecting unit \u2013 This is to allow the shuttlecocks to be ejected in different directions in the opponent court area.\nControlling of the ejecting speed \u2013 This will enable ejecting of shuttlecocks at different speeds.\nAccording to these technical observations, the machine should possess at least 3 DOF to properly adjust the ejecting unit. The proposed arrangement of the required 3 DOF for the machine, is shown in figure 01.\nIn addition to that shuttlecock ejecting speed\nwill be controlled using motors in the ejecting unit.\n74\nAuthorized licensed use limited to: Murdoch University. Downloaded on June 13,2020 at 23:52:11 UTC from IEEE Xplore. Restrictions apply.",
            "(1)\n(2)\nWhere, is the terminal velocity\n(3)\nHere\nand are the coordinates of the ejecting position while and\nrepresents the ejecting velocity and the ejecting angles.\nBy changing the initial parameters different simulation results were obtained for the trajectory path. Simulations were done in the MATLAB software platform. Figure 2\nshows the simulation results and table 01 shows the respective initial parameters\nwhich were selected as suitable values according to the simulation. These results are slightly different from the values represented in the reference [17]\ndue to both being depend on intuitive observations of the actual trajectories of the shuttlecock.\nIV. TESTING THE MECHANISM\nBadminton shuttlecock has a special structure. Therefore, to identify the factors which are related to the\nshuttlecock and shuttlecock trajectory path, a testing apparatus was developed. Major purposes of this testing apparatus are testing the suitability of two roller mechanism, accuracy testing with respect to initial shuttlecock position and determining to which extend that the shuttlecock will be damaged by using two roller & with different initial positions.\nUsing this apparatus following conclusions regarding the mechanism, were obtained.\nIn terms of accuracy, vertical ejecting position of shuttlecock is suitable for\nbasic shots but not for shots require accurate positioning like net-drop.\nHorizontal ejecting position provides fine trajectory path and more horizontal distance than vertical ejecting position.\nShuttlecocks can be ejected horizontally without any significant damage.\nTwo roller mechanism is capable of successfully ejecting the shuttlecocks.\nAccording the test conclusions, the prototype was designed and developed with the two roller mechanism. Furthermore, the initial contact angle between the shuttlecock and the roller was designed as horizontal in the prototype, as the shuttlecock\nwon\u2019t be damaged significantly according to the test results. However\nthe gap between two rollers, the roller material and shape\nare significantly important factors for obtaining these results.\nV. MODELLING THE PROTOTYPE\nThere are three major sections consisted of this badminton shuttlecock feeding machine. They are the dispenser unit, feeding unit and ejecting unit. Dispenser unit is used to store the shuttlecocks and to\nprovide shuttlecocks one by one to the\nfeeding unit. Feeding unit is used to feed the shuttlecocks from the dispenser to\nthe ejecting point and it need to be quick and smooth.\nEjecting unit is designed to eject the shuttlecocks and it utilizes the two roller mechanism. Apart from these three sections another three motors have to be used in order regulate the angle, direction and height of the shuttlecock ejecting point. Then it will have the capability to produce different shots. The arrangement of the required 3-DOF for this, was discussed in the section 02. From now on this arrangement will refer as the shuttle positioning section\nor direction controlling section. Adhering to these technical and other factors a model was developed using SolidWorks software platform. This model of the prototype is shown in Fig. 3.\n75\nAuthorized licensed use limited to: Murdoch University. Downloaded on June 13,2020 at 23:52:11 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_63_0000295_icems.2019.8921961-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000295_icems.2019.8921961-Figure3-1.png",
        "caption": "Fig. 3. Schematic Diagram of Spiral Waterway And 3D Models of the whole machine after adding waterways",
        "texts": [
            " Therefore, the motor needs to add cooling structure to reduce the temperature of the winding part. IV. EDDY CURRENT LOSS CALCULATION According to the structure of the external rotor of the motor, a spiral waterway is designed before the stator and shaft to reduce the temperature rise of the motor. The helical channel area covers the stator core area as much as possible. The schematic diagram of the waterway and the threedimensional model of the whole machine after adding the waterway are shown in Figure 3. After determining the channel structure, the most direct factors affecting the cooling effect of hub motor are the initial temperature of cooling water, inlet speed and flow pattern. Fluid flow can be divided into laminar flow and turbulent flow, with critical Reynolds number Rec= 2320 as the criterion [10]. When Re<Rec, the flow was laminar; when Re>Rec, the flow was turbulent; the particles of turbulent fluid mixed with each other, forming a turbulent flow pattern. The results show that [10] when the cooling structure of water-cooled motor is determined, turbulent flow is better than laminar flow; when the cooling medium is laminar flow, the temperature of each part of the motor decreases with the increase of flow velocity; when the flow pattern of cooling medium is turbulent, the temperature of each part of the motor decreases with the increase of flow velocity, while the temperature of each part of the motor decreases with the increase of flow velocity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000897_j.matpr.2020.04.031-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000897_j.matpr.2020.04.031-Figure7-1.png",
        "caption": "Fig. 7. SRM infrared image showing the signature of the iron losses (and copper losses in wires).",
        "texts": [
            " The stator poles iron losses are very important because the volume of these parts is small compared to the global weight of the SRM. Unlike other classical motors where the measurement of iron losses is easy, the direct measurement of SRM iron losses is nearly impossible. So we have used an infrared camera (Fig. 8) in front of the motor. The SRM is supplied to operate for an hour, the lateral side of the machine is removed rapidly before the decrease of the motor temperature. An infrared image (Fig. 7) is taken just after. This infrared image shows that the stator poles are the most hot part of the motor. This is the signature of the iron losses. The coupling between thermal signature and iron losses is not studied. It should take in consideration the copper losses in winding. The SRM studied here has a power of 2 kW. The total iron losses estimated are very important (near 50W for 1800 rpm), this is due especially to the shape of currents that are not sine. Also to the geometry and the spatial winding repartition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002893_tasc.2021.3089106-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002893_tasc.2021.3089106-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the EDW model.",
        "texts": [
            " 1051-8223 (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. tics of the levitation force and thrust force are analyzed through simulation and experiments. It is verified that the EDW can not only achieve stable levitation in situ, but also provide continuous thrust force. It is expected to provide reference for the follow-up studies of the EDS maglev. As shown in Fig. 1, the permanent magnet EDW is located above a non-magnetic conductive plate. When the EDW rotates and translates, eddy current will be induced in the plate, developing levitation force Fl and thrust force Ft. By comprehensively considering the suspension performance and economy, the NdFeB N45 was selected as the magnet material, and Halbach array which can produce the strongest magnetic field with the smallest number of magnets was applied after optimized simulation. The specific parameters of the permanent magnet EDW used in the experiment are shown in Table I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure9-1.png",
        "caption": "Fig. 9: calculated a) and measured b) mode shape r = 2",
        "texts": [
            " In addition, the end shields and the flange were considered by an elastic support boundary condition indicated by the blue color in the screw threads of the housing. Moreover, this boundary condition allows torsional mode shapes to occur in the simulation. The experimental setup is shown in Fig. 7. Triaxial accelerometers were placed on the outer housing surface of the machine. At 42points (3 rings with 14 points), the machine was excited using a B&K 8206 impact hammer. The software B&K Connect was used to evaluate the stable mode shapes and eigenfrequencies. In Fig. 8 and Fig. 9, the comparison between the measurement and the simulation are displayed. It has to be noted that the measurement was executed with the motor freely mounted by flexible straps (Fig. 7). However, the FE-analysis includes an elastic boundary condition. Likewise, the experimental analysis was performed on the assembled machine including the end shields as well as the rotor. Because of this additional mass, only three mode shapes are identified confidently. Nonetheless, the simulation shows good accordance with the measured eigenfrequencies, as indicated in table III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002383_iros45743.2020.9340703-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002383_iros45743.2020.9340703-Figure2-1.png",
        "caption": "Fig. 2. Schematic representation of the system. Ow is the origin of the world-fixed frame, where a cable of length L is anchored. The other end of the cable is anchored at Ob, which is the origin of a frame attached to the aerial base. A manipulator is attached below the aerial base. A frame with origin in Oe is defined attached to the manipulator end-effector.",
        "texts": [
            " An additional lower-priority task is added as the Cartesian position of the manipulator elbow, to keep the manipulation site free and damp internal redundant motions. The proposed method is validated with numerical simulations and tested on the real hardware. In this section, the main quantities are described together with the details about the model of the system. We suppose that the main carrier, be it a helicopter or a crane, is just deputed to bringing SAM to the location of interest and sustaining its weight. Thus, in our analysis, we consider the system schematically depicted in Fig. 2, composed of an inextensible mass-less cable anchored at a fixed point on the ceiling, a fully-actuated aerial base anchored to the cable, and a 7 DoFs manipulator attached below the platform, not necessarily at its center of mass (CoM). Note that SAM is designed so that the manipulator base is shifted w.r.t. the UAV CoM\u2014see [14] for more details. We define an inertial frame {W } with origin Ow and axes { Xw,Yw,Zw } oriented as in Fig. 2, a frame {B} fixed to the multi-rotor, with origin at its CoM, Ob, which is equivalent to the geometric center of the multi-rotor, and axes { Xb,Yb,Zb } . A reference frame {E } is attached to the end-effector of the arm, with origin Oe and axes { Xe,Ye,Ze } . We define pb = [ xb yb zb ]T \u2208 R3 and pe \u2208 R3as the position vectors w.r.t. {W } of {B} and {E }, respectively. (\u00b7)T indicates the transpose operator. We will indicate with iR j the orientation of frame j w.r.t frame i, where i, j = {w,b,e} indicate {W }, {B}, and {E }, respectively; when i is omitted, it is intended to be i = w"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002988_s11665-021-05948-1-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002988_s11665-021-05948-1-Figure4-1.png",
        "caption": "Fig. 4 Illustration for the formation of different grain morphologies after rapid solidification in metal AM processes: (a) equiaxed cellular grains, (b) equiaxed dendrites and (c) columnar dendrites",
        "texts": [
            " In terms of microstructure, finer grains are found in 3D printed components, while coarser grains are found in conventionally casted components. This is because the higher cooling rate in metal AM processes induced further grain refinement in the microstructure. An article by Flemings (Ref 49) identified six different grain morphologies during the solidification process of directionally solidified alloys. Among them, the following three grain morphologies are commonly reported in 3D printed steel molds: equiaxed cellular grains, equiaxed dendrites and columnar dendrites. Figure 4 illustrates the relationship between thermal gradient and grain morphologies formed after rapid Table 4 A comparison for the maximum relative density of 3D printed steel molds and the corresponding optimized process parameters used to fabricate it Steel mold material Maximum relative density, % Energy density, E \u00bc P vht, J mm 3 Optimized process parameters Laser Power, P, W Scanning Speed, v, mm s 1 Hatch spacing between scan tracks, h, mm Layer thickness, t, mm H13 (Ref 65) 99.20 106.25 170 400 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure3-1.png",
        "caption": "Figure 3. Kinematics analysis of the CZXB method.",
        "texts": [
            " Therefore, with the CZXB method, it is possible to manufacture an hourglass worm with a larger center distance using a machine tool with a smaller center distance. Theoretical analysis of the CZXB method According to the generating theory of the planar enveloping hourglass worm, as shown in Figure 1, the generatrix S(2) rotates around the axis O2-O2 with the angular velocity vector !2, that rotary movement can be regarded as the planar motion of the rigid body. According to the basic theory of rigid body kinematics,18 the kinematics analysis of the CZXB method is shown in Figure 3. The fixed coordinate system g\u00bc [Og; xg, yg, zg] is the original position of the hypothetical tool gear. The position of the coordinate system g in Figure 2 is as follows: the axis xg is parallel to OZ and have the same direction, the axis yg and OX lie in the same line and have the opposite direction, and the axis zg and O2-O2 lie in the same line. The cutting tool is fixed on the rotary table. Generatrix S(2) is the main working surface of the cutting tool. Assume that the generatrix S(2) is a plane",
            " According to formula (1), there are two specific cases: case 1, \u2019 is a constant, variable xO0 and yO0 change together with t, and then O0P translates in plane xgOgyg, which is called translation motion of rigid body; case 2, xO0 and yO0 are two constants, the variable \u2019 changes together with t, and then O0P rotates in the plane xgOgyg, which is called rotation motion of the rigid body. Therefore, the rotary motion of the cutting tool about the axis O2-O2 can be decomposed into the rotational motion B and the translation motions Z and X. It is concluded that the CZXB method is scientific. According to the synthesis and decomposition theorem of velocity, the velocity of any point P is shown in Figure 3(b). Assume that the rotation speed of the plane figures O0T1T2 around the axis O2-O2 is !2; the rotational speed of any point P around the point O0 is !O0P; the translation velocity of point O0 translation along the arc l is vO0; the translation velocity of any point P relative to the point O0 is vO0P. Straight segment OgO0, OgP, and O0P are represented as vector rO0, rP, and rO0P, respectively. Then, the translation velocity of any point P relative to point O2 is written as follows vP \u00bc vO0 \u00fe vO0P \u00f02\u00de where, vP is the instantaneous translation velocity of any point P"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000729_humanoids43949.2019.9035004-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000729_humanoids43949.2019.9035004-Figure3-1.png",
        "caption": "Fig. 3. Left: cross-section schematic of hybrid RPA. Right: a hybrid RPA.",
        "texts": [
            " At 10-\u03bcm misalignment, transmission drops to approximately 78%, at 20 \u03bcm to 38%, and for 50-\u03bcm misalignment, the transmitted ions drop to less than 3%. Given the scaled-down dimensions of the RPAs that can measure cold, dense plasmas, the impact of the misalignment is larger and therefore, aperture alignment is highly desirable to maximize grid transmission. The hybrid RPA, composed of a Computer Numerically Controlled (CNC) machined 316 stainless steel housing and a set of microfabricated silicon electrodes coated with sputtered tungsten (Fig. 3), is our first-generation ion energy distribution sensor. Silicon was selected as electrode material because of its high melting temperature (1687 K), low thermal expansion coefficient (2.6\u00d710\u22126K\u22121), and low electrical resistivity (5 m .cm). Alumina spacers and rails electrically isolate the electrodes from each other and the housing; these components make physical contact to the grids on a few points, far from the region of influence of the ions, to mitigate shorting through surface breakdown and increase device reliability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002935_iemdc47953.2021.9449597-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002935_iemdc47953.2021.9449597-Figure3-1.png",
        "caption": "Fig. 3. Flux contour plots. 2-pole PM motor at nominal motor current.",
        "texts": [
            " The E-core structure decreases the stator and rotor yoke thicknesses compared with the C-core structure. The main radial AMB parameters are shown in Table III. The SPM motor is designed as 160 kW at 30000 r/min. However, the operational speed range starts from 17000 r/min. To minimize the losses, a 2-pole structure and a distributed winding are applied. The outer stator diameter is kept the same as the outer stator diameter of the competing bearingless motor. The composite bandage thickness is determined by structural analysis. Fig. 3 shows flux contour plots of the bearingless unit and the SPM motor in nominal conditions. The laminated magnet segmentation (two pieces per pole) ensures low eddy current losses and structural integrity. For robust construction, an iron pole is used between the magnet poles in both motors. The bearingless motor is designed to minimize the force error angle, force ripple, and torque ripple while minimizing Authors would like to thank Business Finland for funding the project: \u201dEMBER High temperature high-efficiency oil-free heat pump, Decision number 1745/31/2020"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001142_ilt-12-2019-0542-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001142_ilt-12-2019-0542-Figure1-1.png",
        "caption": "Figure 1 Structural representation of conical hybrid bearing",
        "texts": [
            " This paper selects a conical hybrid bearing as the research object, establishes THD model and single quality rigid rotor system dynamics model. Thermal effect on stiffness and damping coefficients, instability speed and minimum oil film thickness are calculated by FEM and FDM, then analyzing the influence of thermal effect on stability and minimum oil film thickness of high-speed conical hybrid bearing. These parameters are compared with literature (Hong et al., 2009) to verify the correctness of theoretical model. A typical deep/shallow pocket conical hybrid bearing is shown in Figure 1. The lubricant is assumed incompressible and some factors are neglected such as rotational inertial forces of fluid, the elastic deformation of journal and bearing shell as well as tilting of axis andmanufacturing errors. When the oil film is expanded into a circular sector along the generatrix, the dimensionless dynamic Reynolds equation is obtained: @ @l l h 3 m @p @l ! 1 1 l sin2a @ @U h 3 m @p @U ! \u00bc 6BM l d 2 l @h @U 1 12BM l d 2 l \u00ab cosw 1 \u00ab0 u sinw cosa z sina h i (1) The whole bearing is assumed to an adiabatic system, which means all heat generated during themovement is taken away by oil flow, so the dimensionless energy equation is: h h 3 6m 1 l2sin2a BM d l 2 @p @U "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002434_ssd49366.2020.9364256-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002434_ssd49366.2020.9364256-Figure10-1.png",
        "caption": "Fig. 10. Flux density distribution (Maxwell 2D).",
        "texts": [
            " Downloaded on May 17,2021 at 04:28:31 UTC from IEEE Xplore. Restrictions apply. Waveforms of stator winding currents and the induced voltage obtained from Ansys Maxwell software of the both models (2D and 3D) are compared in Fig. 8 and Fig. 9. Using Maxwell 2D and Maxwell 3D, the magnetic flux density distribution may be properly compared to evaluate the influence of the both types of the model using ANSYS Maxwell software [3]. Waveforms of the magnetic field calculated in Maxwell 2D and 3D are shown in Fig. 10 and Fig. 11 for time t = 0.2 s and at the speed of 1500 rpm. In Fig. 10 the maximum flux density value is 2.7069 T. It can be seen from Fig. 11 that the flux density value is 2.6465 T. Small differences between the 2D and 3D solutions in the ANSYS Maxwell software calculated of the permanent magnet synchronous motor are primarily caused by influence of the motor end windings and the slot skewing. From the comparison of currents waveforms, electromagnetic torque and flux density distribution it is clear that the differences between results of two dimensional and three dimensional models are relatively small but must be considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001332_s10338-020-00179-8-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001332_s10338-020-00179-8-Figure1-1.png",
        "caption": "Fig. 1. Contact model of EPFGM coating",
        "texts": [
            " The continuous spatial gradation of all material properties of EPFGM coating through its thickness is assumed. In the wear model, the surface deformation and roughness are both considered. Numerical examples are presented to analyze the effect of worn band on the contact pressure and film thickness of ball bearings, and the graded index effect of EPFGM coating is also examined. Coatings and FGMs are widely used in modern machinery to prevent surface damage and enhance lives of critical components. Oil film exists between the rolling ball and EPFGM coating, as shown in Fig. 1a. In point contacts, the contact region is very small in comparison with the macro-surface dimensions; therefore, a half-space assumption is used. The pressure field of oil film between the EPFGM and rolling ball can be determined by using mixed elastohydrodynamic lubrication, and the Reynolds equation can be expressed as [6]: \u2202 \u2202x ( \u03c1 12\u03b7\u2217 h3 f \u2202p \u2202x ) + \u2202 \u2202y ( \u03c1 12\u03b7\u2217 h3 f \u2202p \u2202y ) = U \u2202\u03c1hf \u2202x (1) where p is the oil pressure, hf is the film thickness, \u03b7\u2217 is the effective viscosity, \u03c1 is the mass density, and U is the entrainment velocity to the contact area",
            " q is the normalized ration of stress to strain transfer, and q = \u03c3c\u2212\u03c3m Ec(\u03b5c\u2212\u03b5m) (0 \u2264 q \u2264 \u221e), where \u03c3c and \u03c3m are the stresses in ceramic and metal phases, \u03b5c and \u03b5m are the strains in ceramic and metal phases, respectively. The EPFGM coating can be well modeled with materials of multilayered material systems, where the layers are perfectly bounded with different material properties in each layer. To obtain the displacement and stress fields, the EPFGM coating is divided into L layers, as shown in Fig. 1b. By introducing the Papkovich\u2013Neuber potential functions \u03d5 and \u03c8(\u03c81, \u03c82, \u03c83), the displacements and stresses in the EPFGM layers can be expressed as u (j) i = 1 2Gj [\u03d5(j) ,i + x\u03c8 (j) 1,i + zj\u03c8 (j) 3,i \u2212 (3 \u2212 4\u03bdj)\u03c8 (j) i ] (j = 1 \u2212 L) (8a) \u03c3 (j) ik = \u03d5 (j) ,ik \u2212 2\u03bdj(\u03c8 (j) 1,1 + \u03c8 (j) 3,3)\u03b4ik \u2212 (1 \u2212 2\u03bdj)(\u03c8 (j) i,k + \u03c8 (j) k,i ) + x\u03c8 (j) 1,ik + zj\u03c8 (j) 3,ik (8b) where subscripts i and k denote the x, y, and z coordinates, Gj = Ej/2(1 + \u03bdj) is the shear modulus of the jth layer, Ej is the elastic modulus of the jth layer, \u03bdj is the Poisson\u2019s ratio of the jth layer, \u03b4ik is the Kronecker delta function, and L is the number of divided layers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure1-1.png",
        "caption": "Figure 1. The generating theory of the planar envelopment hourglass worm.",
        "texts": [
            " The straight line or curved surfaces mentioned above is called the generatrix of the hourglass worm. We usually classify hourglass worm gears according to the type of the generatrix. The above conjugate motion is also referred to as the relative motion between the worm blank and the cutting tool. In the typical case, we define the relative motion based on the position and motion relationship between the hourglass worm and its worm gear. The generating theory of the planar enveloping hourglass worm17 is shown in Figure 1. The shortest distance between rotary axis O1-O1 and O2-O2 is theoretical center distance a. The crossing angle between the axis O1-O1 and O2-O2 is usually 90 . The acute angle between the generatrix S(2) and the axis O2-O2 is . db is the diameter of main base circle. The generatrix S(2) rotates around axis O2-O2 at an angular velocity vector !2. Meanwhile, the hourglass worm blank rotates around axis O1-O1 with the angular velocity vector !1. !1/!2 is called transmission ratio, which is kept constant",
            " Therefore, the CZXB method is more suitable for milling or grinding the tooth surface of the hourglass worm. Obviously, the machining center distance at is constantly changing during the machining process and can be much smaller than the theoretical center distance a. Therefore, with the CZXB method, it is possible to manufacture an hourglass worm with a larger center distance using a machine tool with a smaller center distance. Theoretical analysis of the CZXB method According to the generating theory of the planar enveloping hourglass worm, as shown in Figure 1, the generatrix S(2) rotates around the axis O2-O2 with the angular velocity vector !2, that rotary movement can be regarded as the planar motion of the rigid body. According to the basic theory of rigid body kinematics,18 the kinematics analysis of the CZXB method is shown in Figure 3. The fixed coordinate system g\u00bc [Og; xg, yg, zg] is the original position of the hypothetical tool gear. The position of the coordinate system g in Figure 2 is as follows: the axis xg is parallel to OZ and have the same direction, the axis yg and OX lie in the same line and have the opposite direction, and the axis zg and O2-O2 lie in the same line"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000124_012164-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000124_012164-Figure1-1.png",
        "caption": "Figure 1. Nut skin peeling machine",
        "texts": [],
        "surrounding_texts": [
            "The research was carried out in the Computer Laboratory and Design of the Mechanical Engineering Department of Musamus University. The research method used was an experimental method (computer simulation). The strength of construction with varied loads using the 2015 Autodesk inventor professional software. ICROEST IOP Conf. Series: Earth and Environmental Science 343 (2019) 012164 IOP Publishing doi:10.1088/1755-1315/343/1/012164 Free variables: a load of raw materials (nuts) and machine components. Non-independent variables : main stress distribution, displacement, and safety factor. To determine the strength of frame construction loading is given on the Y-axis of 500 N, 600 N, 700 N, 800 N, 900 N, and 1000 N ICROEST IOP Conf. Series: Earth and Environmental Science 343 (2019) 012164 IOP Publishing doi:10.1088/1755-1315/343/1/012164"
        ]
    },
    {
        "image_filename": "designv11_63_0000063_apusncursinrsm.2019.8889073-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000063_apusncursinrsm.2019.8889073-Figure2-1.png",
        "caption": "Fig. 2. Four elements folded in an accordion origami pattern.",
        "texts": [
            " patch is fed by two microstrip lines, one for each mode of operation, with geometrical characteristics L1 = 18.07 mm, W1 = 1.3 mm, L2 = 14.96 mm, L3 = 33.3 mm, and W2 = 0.95 mm. A 0.6 mm wide slot is introduced in the patch as shown in Fig. 1. Note that 1.5 mm thick FR-4 is used as the substrate. The patch\u2019s first mode is fed by port P1 and resonates at 2.33 GHz, while the second mode is fed by port P2 and resonates at 2.91 GHz. All the elements are uniformly spaced by a distance d along the y-axis. The inter-element spacing is uniformly varied by employing an accordion folding technique as shown in Fig. 2 for four of the seven elements. An accordion structure is selected because practical implementation is achieved utilizing a 3D printed modified scissor lift structure, [4]. Moreover, the mutual coupling between any two ports is always less than \u221210 dB for any practically achievable inter-element spacing d. 411978-1-7281-0692-2/19/$31.00 \u00a92019 IEEE AP-S 2019 The capacity of the communication channel is calculated using (1) as provided by Janaswamy in, [5]: C = log2 ( det [ IM + \u03c1 N \u00d7 ZR\u03a8 RZ \u2020 RH uZ \u2020 T\u03a8 TZTH u\u2020 |CTCR|2 ]) (1) C represents the mean channel capacity in bits/s/Hz over 300 channel realizations and \u03c1 (assumed here to be 10 dB) is analogous to the signal-to-noise ratio (SNR)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000712_auteee48671.2019.9033244-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000712_auteee48671.2019.9033244-Figure12-1.png",
        "caption": "FIGURE 12. Temperature distribution for CPU and two tiers of memory with silicon thinning to 100 \u00b5m.",
        "texts": [
            "375 W each. The results show the worst-case hot spot yields 322.1 K with a temperature variance of about 12 K from Fig. 10. Silicon has a thermal conductivity of about 140 W/m-k, hence it has limited heat dissipation capabilities. One approach to improve this is thinning the bulk Si. Fig. 11 shows the proposed CPU configuration with a 100 \u00b5m Si substrate 48 VOLUME 3, 2015 and two tiers of memory. Temperature distribution of this yields hot spots of 315 K and 11 K variation, which can be seen in Fig. 12. One can see that the heat from the hottest blocks in the processor accumulates with the memory above, spreading to some extent into the bulk Si layer. Further improving this lateral heat spreading is expected to result in a decrease in the temperature of the hot spots. Hence, a strategy suggested here is to remove part of the bulk Si layer and replace it with high thermal conductivity synthetic diamond. VII. THERMAL EVALUATION WITH DIAMOND Using the same conditions and assumptions from the previous section, simulations are conducted using diamond layers and further thinning of the bulk Si"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000419_phm-qingdao46334.2019.8942956-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000419_phm-qingdao46334.2019.8942956-Figure4-1.png",
        "caption": "Figure 4. Deployment of test bench",
        "texts": [
            " Architecture of 1D CNN network Rolling bearing dataset (vibration parameters) Dataset segmentation (training / testing) Training dataset Testing dataset Standardized processing Deep learning modelling Training Standardized processing Testing Results 1D CNN LSTM GRU Networks Training process Testing process 2019 Prognostics & System Health Management Conference\u2014Qingdao (PHM-2019 Qingdao) For each deep learning-based model, use standardized training dataset, which consists of a large amount of samples and is split into training part and validation part, to train and validate the model. In order to better validate the reliability of deep learning models, data from several different working conditions is used to train models. And then use the trained model to test the testing dataset. Finally, we get results from three different models for different working conditions. III. EXPERIMENTAL SETUP The whole test bench includes drive motor, load motor, test gearbox, load gearbox, etc., as shown in Fig. 4. The parallel shaft gearbox is divided into two gear sets, which means the gearbox can be decelerated by two stages, and the deceleration effect is more obvious. There are three shafts in the parallel shaft gearbox test bench, namely the input shaft, the intermediate shaft and the output shaft. The input shaft is connected to the output shaft of the planetary gearbox. The intermediate shaft is the uppermost shaft in Fig. 4. The fault parts replaced are the gear and rolling bearing of the intermediate shaft. The output shaft is connected to the rightmost couplings. The parallel shaft gearbox test bench can be loaded with normal gears (or bearings) or fault gears (or bearings) for life prediction experiments. The load gearbox is used to increase the output speed of the test gearbox output and reduce the torque load. There are 13 sensors deployed on the test bench, including motor Z, planet X, planet Z, planet carrier Y, planet carrier Z, parallel middle left Y, parallel middle left Z, parallel box base Z, parallel middle Y, parallel middle Z, parallel input right X, parallel input right Z and parallel box Y, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001197_s0263574720000533-Figure20-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001197_s0263574720000533-Figure20-1.png",
        "caption": "Fig. 20. The illustration of the simulation scenario.",
        "texts": [
            " 2: for each i \u2208 [1, n0] do 3: for each j \u2208 [1, n] do 4: p(i, j) = Qj + t(Qj+1 \u2212 Qj) 5: solve t by the condition that (Qj+1 \u2212 Qj) \u00b7 (p(i, j) \u2212 ci ob) = 0 6: if t < 0 then 7: t = 0 8: end if 9: if t > 1 then 10: t = 1 11: end if 12: d(i, j) = \u2016ci ob \u2212 p(i, j)\u2016 13: ds(i, j) = d(i, j) \u2212 ri \u2212 Rj \u2212 ltl 14: k = 0 15: if ds(i, j) < 0 then 16: k = k + 1 17: elevate degree of the B\u00e9zier backbone curve using Eq. (47) and generate a new free Dimensions Module: 4.3 cm \u00d7 4.3 cm \u00d7 8.65 cm Length (full 16 module robot): 138.4 cm Length (start link): 4.05 cm Length (head link): 4.6 cm Mass Module: 0.2 kg Full 16 module robot: 3.2 kg Joint limit \u00b199 degree clutter environment constructed by three spheres. The structure parameters of the snake robot are given in Table I. The simulation is completed on the dynamics simulation software. The simulation scenario is shown in Fig. 20. The proportional-derivative controller with gravity compensation is adopted to ensure the accuracy of the trajectory tracking in the joint space. For the obstacle avoidance simulation, the positions of the obstacle spheres are known in advance. The radii of the three obstacle spheres, r1, r2, and r3, are all 5 cm. The radii of envelope cylinder of the links R are all 3.04 cm. The distances between sphere 1, sphere 2, and sphere 3 are 6.8, 7.2, and 7.2 cm respectively. The positions of the spheres are [\u22120.281767, 0.035917, 0.319962], [\u22120.116983, 0.00319943, 0.319914], and [\u22120.199326, 0.0195819, 0.469938] m, respectively. The positions of the end point of the links for the base part of the snake robot are [0, \u22120.0405, 0], [0, 0, 0], [0, 0.0865, 0], [\u22120.06412, 0.1446, 0], [\u22120.1282, 0.2026, 0], [\u22120.2143, 0.194, 0], [\u22120.3004, 0.1855, 0], [\u22120.3518, 0.1159, 0], and [\u22120.40322, 0.04637, 0] m. The head of the snake robot moves through the clutter environment along the direction like black arrows in Fig. 20. https://doi.org/10.1017/S0263574720000533 Downloaded from https://www.cambridge.org/core. University of Exeter, on 03 Jul 2020 at 02:30:12, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. The specific process of obstacle avoidance is shown in Fig. 21. The initial state of the snake robot is shown in Fig. 21(a). Then the snake robot head approached the spheres in Fig. 21(e). The three intermediate states for approaching the spheres are designed based on the B\u00e9zier backbone curve as shown in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001869_s11223-020-00217-3-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001869_s11223-020-00217-3-Figure11-1.png",
        "caption": "Fig. 11. Critical elements found according to the analysis results of the transient finite element model of the wheel.",
        "texts": [
            " As mentioned above, the transient strain histories reached their respective steady states after approximately 2 s, after which their amplitudes were practically unchanged with time, with the fatigue damages of the wheel mainly caused by the steady parts of the transient radial strain histories. Thus, only the steady radial strain histories of the critical elements were used to predict the fatigue life of the wheel. It is important to identify the critical elements in the finite element model of the wheel that are responsible for predicting the wheel fatigue life, because their fatigue lives are the shortest. Figure 11 shows the critical elements found according to the steady radial strain histories obtained from the transient finite element model of the wheel. Figures 12 and 13 show the radial strain histories of the critical elements in the left and right part of the wheel model, respectively, and their representative steady parts are shown in Figs. 14 and 15, with each of them having 801 sampling points, which are the last 801 points and about 20 cycles of the total 1601 transient response points corresponding to the time length of 5",
            " As can be seen, there are 20 large hysteresis curves which are basically identical and whose local stress-versus-strain characteristics are significantly non-linear and include 1 small loop. The fatigue damage Da caused by these 21 cycles (loops) is 5.75884e-04. By contrast, the damage caused by the small loop is only 6.50438e-19, which means that its contribution to Da is minimal. Thus, Da can be considered as caused only by the 20 large hysteresis loops or cycles. This assumption is also valid for all other critical elements shown in Fig. 11. Presuming that L f is the predicted fatigue life or the predicted constant amplitude of loading cycles the element can withstand before an engineering crack is formed, the formula for L f is as follows: L D f a 20 . (6) Each critical element of fatigue life L f is also shown in Table 6, where the predicted fatigue lives are quite close to each other. Generally speaking, the shortest life is the most important, and is usually taken as the predicted structure life. Moreover, as can be seen in Table 6, the element 1R has the shortest life of 34,689 cycles, which was taken as the predicted fatigue life of the wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003200_s00419-021-01997-z-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003200_s00419-021-01997-z-Figure10-1.png",
        "caption": "Fig. 10 Scissor deployable structure composed of two SLEs",
        "texts": [
            " In addition, it should be noted that this model associated with the formula (34) only corresponds a simple cantilever beam structure, while the planar 2\u00d72 deployable structure analyzed here is complex structure including rigid body displacement and bar deformation as well as their coupling. Therefore, when the bar length increases to a certain value that is 0.1 m observed in Fig. 9a, the decrease in the frequency becomes slower due to the interaction between different members. Similarly, it can be observed from formula (34) that the change of natural frequency is directly proportional to the rigidity of the bar, and the change trend is basically consistent, which can be confirmed from Fig. 9b. 4.3 Dynamics of deployable structure composed of two SLEs Figure 10 shows a 1\u00d72 deployable structure composed of two scissor elements, where point C has a degree of freedom along the Y-axis and bears a downward external load, P 100 N. The translational degrees of freedoms at point A are constrained, and \u03d51 and \u03d52 are the initial unfolding angles of bar 2 and bar 1 relative to the X-axis. The main structural parameters of the bar used for numerical analysis are as follows: The bar length is 2 m, and the cross-section is 0.02 \u00d7 0.04 m2. The initial deployment angle and the material of the bar are 60\u25e6 and aluminum alloy, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000834_crc.2019.00024-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000834_crc.2019.00024-Figure8-1.png",
        "caption": "Fig. 8. Control of rotary motions of the robot shoulder by using the body: (a) front body of the robot and (b) side view of stopper on the robot front",
        "texts": [
            " As configured for this study, the length of the upper arm link [from Joint 2 (shoulder) to Joint 4 (elbow)] was l2 and the length of the forearm link (from Joint 4 to Joint 6) was l4C. The manipulator joint angles were \u221290 deg \u2264 \u03c62 \u2264 +90 deg and 0 \u2264 \u03c63\u2264 +100 deg (Fig. 3). The right and left hands each had two fingers (Joint 6) to hold items (Fig. 6). Touch sensors were set on both forearm links, each of which were covered by bumpers (Fig. 7). These touch sensors were used when the robot front wheels climb a step (see Section III). The robot had a stopper mounted on its front of its body (Fig. 8 (a)). The stopper limited the passive rotational travel of the manipulators during wheelchair pushing (Fig. 8 (b)) and enabled the robot to imitate the operation of a human pushing an object (Fig. 9). The robot did not need to exert 76 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 06:18:12 UTC from IEEE Xplore. Restrictions apply. force around the shoulder axes, because the upper arms pushed against the chest (i.e., the stopper) when the robot pushed the wheelchair. The wheelchair (NOVA Integral-ME) in this study was a commercially available electric model with rearwheel drive (TRD-1, Acritech Co",
            " <2> The robot stops, but the wheelchair continues forward until its accelerometer system detects wheelchair inclination, which means the system is ready to place the front wheels of the wheelchair on the step. <3> Initially, the wheelchair center of mass is in front of the contact point between the rear wheels and the ground. When the mass shifts to behind the contact point as the wheelchair tilt increases, the stopper of the robot limits the passive rotation about Joint 2 (Fig. 6) and prevents the wheelchair from tipping over backward (Fig. 8 (b)). <4> Both vehicles move forward, and the front wheels of the wheelchair are positioned on the step. Stage 2 <5> Both vehicles continue to move forward. <6> The back wheels of the wheelchair come into contact with the step. <7> The robot continues to push the wheelchair so that the rear wheels of the wheelchair climb up onto the step. The 78 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 06:18:12 UTC from IEEE Xplore. Restrictions apply. robot supports the wheelchair during this process to prevent the wheelchair from tipping over backward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure8-1.png",
        "caption": "Fig. 8. Schema of the spring based alternative model.",
        "texts": [
            " This can be observed with the MATRIX27 user-defined matrix element for type B roller. The components of this matrix are identical to the type A roller matrix with the exception of k12, k14, k21, k23, k32, k34, k41 and k43, which sign changes. I. Mart\u0301In et al. Tribology International 161 (2021) 107098 As mentioned before, instead of using the MATRIX27 element to simulate the wire-roller-wire set, a simple compression-only spring mechanism that simulates the deformation of the set could be used. As shown in Fig. 8, the spring is able to represent the compatibility of deformations of the set under different axial or radial external displacements. However, this alternative model does not include friction forces, so wire twisting and its eventual consequences are not considered. The stiffness of the equivalent spring (kcbn) is obtained from an equation based on geometrical relationships of k1, k2 and k3 (Fig. 9a). First, it is necessary to combine the stiffness of k1 and k2 into a kp projecting them on the roller-wire contact normal line defined by the initial contact angle \u03b10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure53.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure53.3-1.png",
        "caption": "Fig. 53.3 a Reaction chamber, b sputum collection container, c leak-proof transfer of sample",
        "texts": [
            " The two sections of processing unit are covered by vastly distinct lids equipped with indicators and electronic locks to encourage the right action and prevent the wrong ones. Clear indicators for device on/off, Bluetooth connectivity and battery life have been provided for effective and unambiguous indication (Fig. 53.1) Disposables A set of two single-use disposables are to be used in succession and to hold the sample in complete containment. The first, called sputum collection and preparation container (SCC), is used to collect a sputum sample and to prepare it for steps to follow (Fig. 53.3b). The second, called the reaction chamber (RC), holds the sample in processing unit, while creating a physical interface between sample and machine, without hampering the test or sample within (Fig. 53.3a). Both these disposables hold pre-measured reagents to be used before and during test procedure. Provision of self-healing rubber and incorporation of vacuum vials 648 G. Garg et al. ensure perfectly leak-proof transfer of measured amounts of sample solution from one to the other, when the process demands it (Fig. 53.3c). The design incorporates mechanically driven mechanisms internally, such as to release reagent, to filter out sputum from spit and to transfer solution fromcentrifuge friendly vial to PCR friendly microwells. User Interface The complete process is navigated and controlled by a dedicated app, called the MycoKitApp. The app tracks the ongoing test through checkpoints and error announcement and works as a dashboard for quick access to test results, patient history, system status, platform to give feedback and alert for errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000409_0954406219896815-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000409_0954406219896815-Figure5-1.png",
        "caption": "Figure 5. Data processing for machining simulation.",
        "texts": [
            " In the simulation, the tool outline is represented by a finger cone. The three-dimensional model of the finger cone is constructed in VERICUT based on the shape and geometry of the worm tooth groove. The parameters of the worm tooth are shown in Table 1. Machining simulation focuses on the relative motions between the tool and the worm blank. Processing simulation ignores many details of the actual production in this paper. Taking the finishing process of the hourglass worm teeth as an example, data processing for machining simulation is shown in Figure 5. The interpolation plane is XOZ. The worm blank rotates with the C axis. The tool makes a rotational motion about its axis, which is the main motion during the cutting process. The tool is fixed on the B-axis rotary table and rotates with the B-axis turret. At the same time, the B-axis turret can be translated along the X and Z axes. The translational trajectory of the B-axis rotation center is the arc l in finishing step. The radius of the arc l is rc1. ps is the start point of the cutting motion. pe is the end point of the cutting motion. When the B-axis rotation center is at the ps point, the initial state of the process system is represented by [xs, zs, cs, bs]. When the center of rotation of the B-axis is at point pe, the state of the process system is represented by [xs, zs, cs, bs]. According to Table 1, Figure 5, and the tool setting operation, the above coordinate values are easy to determine. Based on the Fanuc oi CNC system, the core G code for simulation are shown below: G00 Xxs Zzs Ccs Bbs; G18; G02 Xxe Zze Cce Bbe Rrc1; M30; The machining model of the hourglass worm is shown in Figure 6. Since the tool refers to a finger tapered tool, the virtual machining model is a finger tapered surface enveloping hourglass worm according to the meshing principle. As shown in Figure 6, the virtual model is a right-handed multi-threads hourglass worm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002338_012093-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002338_012093-Figure4-1.png",
        "caption": "Figure 4. Real schema and equivalent force in case where concentrate force acting on bar.",
        "texts": [],
        "surrounding_texts": [
            "The bar can also be placed obliquely, thus forming an angle with the horizontal axis. In this case the axis from the first node of the element to the second node is called the local axis. The dependence 10th EASN 2020 IOP Conf. Series: Materials Science and Engineering 1024 (2021) 012093 IOP Publishing doi:10.1088/1757-899X/1024/1/012093 between forces and displacements derived in relation to these axes, called rigidity matrix in local coordinates. We consider for each node three degrees of freedom and the general case of the element with three degrees of freedom. The forces at each node according to the local axes are normal force, shear force and bending moment. They are located in both element nodes and they form the vector of forces in local coordinates. The same is done for generalized displacement of element nodes as it showed in equation (5). \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f = \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f (5) The element may have loads from the outside, which we equivalent them in the nodes according to the templates (figure 5a, 5b and 5c). These forces are called equivalent forces. They form the vector of equivalent forces, which must be identified to be deducted at the end of the calculations, because it is something imaginary and serves for a specific purpose not for all purposes. The same we can do for displacement [4]. The general equilibrium of the element is expressed in this matrix form, equation (6). \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f = \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f + \ud835\udc5f\ud835\udc5f\ud835\udc3f\ud835\udc3f.\ud835\udc4e\ud835\udc4e\ud835\udc52\ud835\udc52 = \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f \u2212 \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f.\ud835\udc4e\ud835\udc4e\ud835\udc52\ud835\udc52 \u27f9 \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f + \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f.\ud835\udc4e\ud835\udc4e\ud835\udc52\ud835\udc52 = \ud835\udc60\ud835\udc60\ud835\udc3f\ud835\udc3f\ufffd = \ud835\udc58\ud835\udc58\ud835\udc3f\ud835\udc3f \u2219 \ud835\udc51\ud835\udc51\ud835\udc3f\ud835\udc3f (6) Rigidity matrix in the local coordinates for the general case according to the theory of structure [2] is showed in figure 8."
        ]
    },
    {
        "image_filename": "designv11_63_0000189_ecce.2019.8912698-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000189_ecce.2019.8912698-Figure4-1.png",
        "caption": "Fig. 4. Velocity field vectors inside the slotted air gap at Ta=31166",
        "texts": [
            " Software Validation As a validation of the model, the results obtained numerically are compared to those calculated from empirical correlations of the literature. The difference between numerical and analytical results is reasonable and leads to accept that CFD method can be used to estimate windage losses. IV. ANALYSIS AND RESULTS A. Velocity field The rotor slot leads the airflow to generate vortices within the slot. The airflow is assumed to have the same velocity of the boundary at fluid\u2013solid boundaries. The flow is rotating at the same velocity near the rotor and is fixed near the stator. In Fig. 3 and Fig. 4 the velocity field is represented, from the stationary frame, for the rotating configuration at \u2126=20000 rpm (Ta=31166). The airflow is initially regular in the smooth part. At the slot, the change in thickness of the air gap makes the streamlines deviate towards the bottom of the slot with a swirling motion. A stagnation zone with a weak airflow recirculation can be detected and Taylor-Couette structures can as well be observed inside the slot. As the rotation speed increases, the turbulence level inside the slot increases accordingly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002175_978-981-15-7711-6_18-Figure18-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002175_978-981-15-7711-6_18-Figure18-1.png",
        "caption": "Fig. 18 Z displacement plot after spring back",
        "texts": [
            " The normal distance between the root plane and tip plane is 5 mm with a tolerance of \u00b11 mm. From Fig. 16 it can be said that maximum thinning is less than 0.5% which is negligible (well below the 10% limit) From Fig. 17, it can be observed that the portion of leading edge near root has deformed more compared to the portion of leading edge near tip. This is because of relatively greater spring back observed at that portion. As discussed previously, the spring back is not uniform and the compensation should be in accordance with the spring back. From Fig. 18 it can be seen that the leading edge and tip nearly lie on the same plane with acceptable deviation from the required. The portion at the centre of the leading edge seems to be deviating but the deviation is acceptable or can be sort out manually. Also, the third criterion that is the normal distance between the root plane and tip plane is also satisfied. From the above discussion, it can be concluded that the suggested forming process and die design are acceptable in the manufacturing of wing panel and the process can be controlled better compared to the previously discussed one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002142_jmems.2020.3047774-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002142_jmems.2020.3047774-Figure3-1.png",
        "caption": "Fig. 3. Models and parameters of shape-recognizable origami sheet device: (A) Definition of hinge positions. (B) Detailed parameters regarding the length between hinges.",
        "texts": [
            " The coordinates of each measurement point are obtained by measuring the angle of the two hinges, which allows for the measurement of the surface geometry of the object. Figure 1 (B)(i) shows the device deformation if only stretching occurred. On the other hand, Fig. 1 (B)(ii) shows the case of simultaneous stretching and bending. By determining the coordinates of each measurement point, this proposed system can identify deformations. The entire shape of the shape-recognizable origami sheet device is reconstructed by the geometric condition of the device using lines and joints (Fig.3 (A)), where the length between the hinge parts (Ln), the bending angle of the hinge parts (\u03b8n) and the angle of the hinge parts from the horizontal axis (\u03d5n) are defined. The subscript n means the order of the hinge parts from the edge of the device. The angles of the valley fold and the mountain fold are defined as positive and negative values, respectively. Considering the detailed deformation of the hinge part due to bending, we divide Ln of each hinge part into three parts, Ln_HingeB , Ln_Rigid , and Ln_HingeF (Fig.3 (B)). Ln_Rigid is the length of the rigid part determined by the device design. Ln_HingeB and Ln_HingeF are calculated by assuming that the deformation of each hinge part is an arc. Thus, Ln_HingeB and Authorized licensed use limited to: University of Cape Town. Downloaded on May 17,2021 at 21:08:49 UTC from IEEE Xplore. Restrictions apply. Ln_HingeF are expressed below, Ln_HingeB = tan \u03b8n\u22121 2 \u00d7 ( L Hinge \u03b8n\u22121 + T \u2212 t 2 ) (2) Ln_HingeF = tan \u03b8n 2 \u00d7 ( L Hinge \u03b8n + T \u2212 t 2 ) (3) where the thickness of hinge parts is t and the thickness of rigid parts is T . Note that when \u03b8 is 0, the Ln_HingeB and Ln_HingeF are L Hinge /2. By using these equations (2) and (3), the coordinates of each joint in the two-dimensional x-y plane, pn, (Fig. 3 (A)) are calculated as follows: pn = Ln (cos \u03d5n, sin \u03d5n) + pn\u22121 (n \u2265 1) (4) \u03d5n = \u03d5n\u22121 + \u03b8n (n \u2265 1) (5) \u03d50 = \u03b80 (6) The position of each joint can be determined by solving the recursion formula above with initial values. By using these calculations, the entire shape of the origami sheet device is reconstructed. Two types of origami sheet devices were designed for our experiments: (i) the SWCNTs strain sensor device for measuring the characteristics (Fig. 4) and (ii) the shaperecognizable origami sheet device for demonstrating shape and motion monitoring (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000440_s11071-019-05457-w-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000440_s11071-019-05457-w-Figure1-1.png",
        "caption": "Fig. 1 The mechanical soft impacting system",
        "texts": [
            " We have fabricated the proposed circuit and have confirmed that the dynamical phenomena observed in the mechanical systems are also observed in the model electronic circuit. The advantage offered by the electrical analog is further illustrated by experimentally observing the evolution of the chaotic attractor that occurs at the grazing condition, as the stiffness ratio is varied. We have also experimentally obtained the bifurcation diagrams with different values of stiffness ratio [19,32]. A schematic diagram of the mechanical system under study is depicted in Fig. 1. It is basically a forced damped oscillator with a massless compliant wall that the mass M can impact with. By increasing the stiffness ratio k2/k1, one can bring the behavior close to that of a system with instantaneous or hard impacts. The electronic analog of this system proposed in this paper is shown in Fig. 2a. The system consists of an LCR circuit with an input voltage Vin which is a sine wave of amplitude Vamp with frequency f . The system has an analog switch S, controlled by a comparator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002819_asjc.2566-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002819_asjc.2566-Figure1-1.png",
        "caption": "FIGURE 1 Typical fuzzy membership functions",
        "texts": [
            " The Laplacian matrix L= [lij] RN\u00d7N of followers is defined as lij= \u2212 \u03b1ij if i\u2260 j and lij = PN k\u2260i,k=1 \u03b1ij if i= j, then the following lemma is obtained. Lemma 1. [55]: Let the directed graph be connected, then the matrix L=L+ diag \u03b110 \u03b120 \u2026 \u03b1N0\u00bd \u00f0 \u00de is positive definite with \u03b1i0 in the (i, i) diagonal position, and \u03b1i0 = 1 if the leader can transmit the state information to followers, whereas \u03b1i0 = 0 if otherwise. 2.2 | The type-2 FLS Comparing with type-1 FLS, the interval type-2 FLS enhances the ability to describe and deal with uncertainties. A typical type-1 ordinary fuzzy membership function is shown the Figure 1A, and the type-2 fuzzy membership function as description of Figure 1B is derived by blurring the boundaries of type-1 ordinary fuzzy membership function. As it can be seen from Figure 1, the membership grade of the input x* has one certain value \u03bcA(x*) in Figure 1A. Different from the type-1 Gaussian membership function, the membership function of type-2 has a fixed standard deviation \u03c3 with uncertain values in [m1, m2] in Figure 1B, the membership grade of the input variable x* between the lower bound membership function \u03bc A x \u00f0 \u00de and the upper bounded membership function \u03bcA x \u00f0 \u00de , it is known that the fuzzy sets of Type-2 have blurred boundaries, which means that Type-2 has uncertainties properties, so it can reduce the effect of uncertainties of systems. In this research, the interval type-2 FLS is adopted to approximate the unknown nonlinear functions. The function blocks of the FLS mainly include fuzzification, rule base, fuzzy inference engine, and defuzzification mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003241_ldia49489.2021.9505999-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003241_ldia49489.2021.9505999-Figure2-1.png",
        "caption": "Fig. 2. Structure of proposed DSLRPMG",
        "texts": [
            " STRUCTURE AND OPERATION PRINCIPLE DSLRPMG A new type of DSLRPMG applied in the novel WWCEC system is proposed in this section. The DSLRPMG includes an outer stator, a mover, an inner stator, linear unit windings, rotary unit windings, linear unit permanent magnets and rotary unit permanent magnets. The permanent magnets and coils are assembled in two fixed modules (the outer stator and the inner stator) to enhance the mechanical structure strength and reliability and improve the heat dissipation of this generator. The specific structure of the DSLRPMG is shown in Fig. 2. Fig. 3 shows the equivalent magnetic circuit of the DSRLPMG at different positions. Taking the phase A of rotary unit as an example, when the mover is at position 1, the magnetic flux passes through four winding coils of phase A from the inner diameter to the outer diameter duo to the existence of permanent magnet. At this time, the flux linkage of phase A (\u03c8RA) has the maximum positive value. When the mover is at position 2, the \u03c8A has the maximum negative value. The operation process of the linear unit is similar with that of the rotary unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000068_edpe.2019.8883869-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000068_edpe.2019.8883869-Figure2-1.png",
        "caption": "FIGURE 2. No load magnetic flux line of motors. (a) 48 slots/ 44 poles. (b) 48 slots/ 46 poles. (c) 48 slots/ 50 poles. (d) 48 slots/ 52 poles.",
        "texts": [
            " These four motors are all PMmotor with the same power (20kW) and structure parameters, just as shown in Table II. They are all surface magnet motors with open type stator slot and should also meet the conditions as follows: 1) Fundamental magnetic field produced by stator should be same; 2) Fundamental magnetic field produced by rotor should be same; no-load fundamental wave of EMF should be same; 3) Average value of electromagnetic torque at rated current should be same. III. COMPARISION AND ANALYSIS OF ELECTROMAGNETIC CHARACTERISTIC OF DIFFERENT MOTOR SCHEMES Fig.2 shows the structure and no-load magnetic field distribution of four motors. Fig.3 describes magnetic flux density distribution at no load of four motors, and Table III shows the harmonic components of flux density in Fig.3. For motor with 44 poles, its main wave (order 22) is nearly 0.9T, while other components are quite little. Fig.4 is the magnetic flux density harmonic distribution at rated load of four motors, and the harmonic components of VOLUME 4, 2016 8537 flux in Fig.4 is shown in Table III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001378_s11668-020-00961-3-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001378_s11668-020-00961-3-Figure3-1.png",
        "caption": "Fig. 3 Sliding composite leaf spring",
        "texts": [
            " According to the type of connection with the vehicle body, longitudinal composite leaf springs can be divided into eye-end leaf springs and sliding leaf springs. The common structure of the eye-end leaf spring is shown in Fig. 2. Both ends of the leaf spring are metal eye joints, the leaf spring body and the eye joints are connected by bolts. Two metal plates are attached to the middle of the leaf spring body for eliminating the stress concentration when the U-bolt is clamped. The structure of the sliding leaf spring is shown in Fig. 3. The middle structure of the sliding leaf spring is basically as same as that of the eye-end leaf spring. On both ends of the leaf spring are gaskets and hooks, usually made of metal. The gaskets are mainly responsible for the force transmission between the spring body and the support, and the hooks can prevent the leaf spring from leaving the suspension. The spring body, gaskets and hooks are connected by bolts. For spring body design, the main concerns are its material utilization, fiber volume content and manufacturability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001147_j.jmmm.2020.167119-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001147_j.jmmm.2020.167119-Figure6-1.png",
        "caption": "Fig. 6. Winding diagram of 6-phase IM.",
        "texts": [
            " In this model, a different number of phases combination is used. Polyphase slip-ring IM was used for simulation, the machine specifications are given in Table 1. 7. Result and discussions The algorithm of the proposed approach, shown in Fig. 5, is utilized to obtain the magnetic field analysis, mutual inductance and equivalent circuit, and instantaneous torque. The presented results will be performed at 6 phases in the stator and 6 phases in the rotor. The winding distribution is presented in Fig. 6. The machine winding configuration for one pole is illustrated in Fig. 7. By using FEMM software the magnetic field of the polyphase IM can be deduced as shown in Fig. 8. Also, the mesh was applied by the FEMM preprocessor for breaking machine geometry into several small triangular elements. The number of obtained nodes is 145,716 and the number of elements is 291230. The air gap flux density between analytical formulation using Eq. (12) and the FEMM is shown in Fig. 9. In this figure, there is a slight difference between these results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001345_s11668-020-00932-8-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001345_s11668-020-00932-8-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of mini-traction test machine",
        "texts": [
            " Many researchers have studied on the friction characteristics and lubrication mechanism of some commercial lubricating greases, but studies of the friction mechanism of lubricating greases remain rare, therefore, the friction characteristics and friction mechanism of two types of lithium grease widely used in engineering practice were studied using a mini-traction machine and a microscope. The research results in this paper can provide design and manufacturing basis for lubricating grease manufacturers and users and provide a theoretical basis for the development of friction mechanism and lubrication mechanisms in lubricating grease. Test Machine and Test Principle The test was carried out on the mini-traction machine (Fig. 1). The steel ball and steel disc are made of 52100 bearing steel. The diameters of the ball and disc are 19.05 and 46 mm, respectively. The roughness of the surface of the ball and the surface of disc is 0.01 lm. The contact between the ball and the disc is used to simulate the contact between the balls and the raceway in the bearing. The linear velocities of the ball and the disc at the contact point U1 and U2 can be adjusted by two servo-motors to obtain various rolling velocities U (U = (U1 ? U2)/2) and sliding velocities DU (DU = U1 U2), which can realise continuous changes in slide-to-roll ratio S (S = DU/U) from 50 to 50%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002739_j.matpr.2021.04.018-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002739_j.matpr.2021.04.018-Figure2-1.png",
        "caption": "Fig. 2. 3D Model of Knuckle Joint.",
        "texts": [
            " It is defined by eight nodes having three degrees of freedom at each node: translations in the nodal , y, and z directions. The element has plasticity, hyper elasticity, stress stiffening, creep, large deflection, and large strain capabilities. It also has mixed formulation capability for simulating deformations of nearly incompressible elastoplastic materials, and fully incompressible hyperplastic materials. Various element technologies such as B-bar, uniformly reduced integration, and enhanced strains are supported. The meshed image is given in Fig. 2. The average weight of the vehicle is around 1150 kgs. The design/load calculations are based on this majorly. A worst case load acting on the Knuckle combining all the three load cases was taken for analysis considering conservative approach. These loads when act along with the self-weight makes it more conservative to be considered for the static analysis, ensuring safe design. Bump load case (3 g) = 3*9.81*Kerb weight = 33844.5 N (four wheels) Braking load case (2 g) = 2*9.81*Kerb weight = 22563 N (four wheels) Braking load case (1 g) = 1*9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002670_jestpe.2021.3071923-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002670_jestpe.2021.3071923-Figure1-1.png",
        "caption": "Fig. 1. Topology of the proposed HEBFM.",
        "texts": [
            " Therefore, the extension of flux control regulation and speed range is achieved. In this paper, the operating principles based on winding factor modulation for the proposed HEBFM are presented. Besides, simulation based on finite element method (FEM) and numerical analysis is conducted to verify the effectiveness of the winding factor modulation in flux control applications [23-25]. Furthermore, design and optimization information is provided. Last but not least, a prototype of the proposed HEBFM is manufactured and tested to validate the performance. Fig. 1 shows the 3-D schematic diagram of the investigated HEBFM with labels for each coil. Its main design parameters are shown in Table I. It is worth noting that the arrows drawn on the PM segments indicate the radial magnetization directions. To explain the operating principles of the proposed machine, flux path analysis is carried out. Without field excitation, the HEBFM works like a typical biased flux machine (BFM). For clarity, the flux path of this BFM in the minimum module are presented in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001701_0954406220964512-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001701_0954406220964512-Figure12-1.png",
        "caption": "Figure 12. Line diagram of the 5-link spatial manipulator.",
        "texts": [
            " Case study 3: Trajectory tracking during a medical surgery using ANFIS In this section, a case study of surgical manipulator is taken, which is a 5-DOFs spatial manipulator. Proximal variant of D-H parameter method is used for the kinematic study of this manipulator, whose prototype was developed at CSIR-CSIO, Chandigarh. It is a patient-side manipulator, which is used to track the movements of a surgeon-side manipulator during a robot-assisted surgery. Figure 11 shows the physical prototype of the 5- DOFs spatial manipulator and Figure 12 shows the line diagram with DH frame assignment. The DH parameters of this manipulator are calculated and tabulated in Table 5. Incorporating the obtained DH parameters from Table 5, the tool-tip to base transformation matrix is given as, O 6 A \u00bc6O 1 A 11 2A 22 3A 33 4A 44 5A 55 6A 6 \u00bc 0:94 0:34 0 0 0 1 0:34 0 0:94 0 0 0 \u20d3 \u20d3 \u20d3 \u20d3 \u20d3 1159:58 0 1259:07 1 2 66664 3 77775 (33) For more details on the complete kinematic model of this manipulator, kindly refer.35,36 To present a realistic case-study, the above 5-DOFs manipulator is considered to track the projected trajectory of the tool in a minimal invasive surgery (MIS)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000880_j.mechmachtheory.2020.103900-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000880_j.mechmachtheory.2020.103900-Figure1-1.png",
        "caption": "Fig. 1. Two-link robot arm driven by six muscle-like actuators, m 1 , m 2 , . . . , m 6 . The antagonistic pairs of monoarticular actuators, ( m 1 , m 2 ) and ( m 3 , m 4 ), rotate J 1 and J 2 . The antagonistic pair of biarticular actuators, ( m 5 , m 6 ), drives both joints. It is assumed that the moment arms ( r 1 for J 1 and r 2 for J 2 ) are constant.",
        "texts": [
            " In Section 5 , we reformulate the force distribution problem based on the decomposition and provide a particular solution to the distribution problem. In Section 6 , numerical simulations are conducted to validate the proposed method. Section 7 discusses the physical meaning of the subspaces and the applicability of the proposed method to another class of musculoskeletal robot arms. Finally, Section 8 concludes the paper. 2. Modeling 2.1. Modeling of the robot arm This paper deals with the special class of musculoskeletal robot arms depicted in Fig. 1 [28] . The robot arm has two rotational joints, J 1 and J 2 , which are driven by six muscle-like actuators, m 1 , m 2 , . . . , m 6 . The antagonistic pairs of monoarticular actuators, ( m 1 , m 2 ) and ( m 3 , m 4 ), rotate J 1 and J 2 . The antagonistic pair of biarticular actuators, ( m 5 , m 6 ), drives both joints. Let O \u2212x 0 y 0 z 0 be a fixed coordinate system, where the z 0 -axis coincides with the joint axis of J 1 . The configuration of the robot arm can be represented by the joint angle vector q = [ \u03b81 , \u03b82 ] T \u2208 R 2 , where \u03b81 denotes the angle between the x 0 -axis and the first link, \u03b82 denotes the angle between the first link and the second link, and R denotes the set of real numbers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001663_icuas48674.2020.9214003-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001663_icuas48674.2020.9214003-Figure2-1.png",
        "caption": "Fig. 2. Quadrotor orientation.",
        "texts": [
            " The equations that represent a generic UAV [12] can be described as follows: p\u0307np\u0307e p\u0307d = c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8  uv w  (1) u\u0307v\u0307 w\u0307  = r v \u2212 q wpw \u2212 r u q u\u2212 p v + g \u2212s\u03b8c\u03b8 s\u03c6 c\u03b8 c\u03c6 \u2212 1 m 00 f  (2) \u03c6\u0307\u03b8\u0307 \u03c8\u0307  = 1 s\u03c6 s\u03b8/c\u03b8 \u2212c\u03c6 s\u03b8/c\u03b8 0 c\u03c6 \u2212s\u03c6 0 s\u03c6/c\u03b8 c\u03c6/c\u03b8  pq r  (3) p\u0307q\u0307 r\u0307  = I\u22121 pq r \u00d7 I pq r + \u03c4\u03c6\u03c4\u03b8 \u03c4\u03c8  (4) where cx = cos(x) and sx = sin(x). m, g and I denote the total mass of the UAV, the acceleration of gravity and the inertia matrix, respectively. \u03c4 = [ \u03c4\u03c6, \u03c4\u03b8, \u03c4\u03c8 ]T denotes the total external moments applied and is obtained from (7). By using the UAV\u2019s geometry represented in Fig. 1, the total thrust f can be computed as f1 + f2 + f3 + f4. The torque values \u03c41 and \u03c42 are computed according to the axes\u2019 directions in Fig. 2. This corresponds to (5)-(7), where d denotes the UAV diameter and mi denotes the torque generated by each propeller. Aerodynamic forces are considered perturbations of the system, since they have small influence compared to the ones produced by the motors and the mass of the UAV. \u03c4\u03c61\u03c4\u03b81 \u03c4\u03c81  = d 2  \u2212 \u221a 2 2 \u221a 2 2 \u221a 2 2 \u2212 \u221a 2 2\u221a 2 2 \u221a 2 2 \u2212 \u221a 2 2 \u2212 \u221a 2 2 0 0 0 0   f1 f2 f3 f4  (5) \u03c4\u03c62\u03c4\u03b82 \u03c4\u03c82  =  0 0 \u2212m1 +m2 \u2212m3 +m4  (6) \u03c4\u03c6\u03c4\u03b8 \u03c4\u03c8  = \u03c4\u03c61\u03c4\u03b81 \u03c4\u03c81 + \u03c4\u03c62\u03c4\u03b82 \u03c4\u03c82  (7) UAV\u2019s mass was found by using a scale and is m = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000163_isncc.2019.8909193-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000163_isncc.2019.8909193-Figure2-1.png",
        "caption": "FIGURE 2. The conceptual Rings overlay.",
        "texts": [
            " A brief, early description of RIP can be found in [7]. RIP forms the Rings overlay based on localized operations. This implies that only one-hop neighbor information is exchanged among nodes. RIP dynamically discover the rings at setup time. It then adaptively allocates the MPs at run time. The allocation of the mobile probes is specific to each application and should accomplish the application\u2019s goals. Fig. 1 depicts a possible outcome of RIP when applied to an arbitrary physical network topology, whereas Fig. 2 provides an abstracted view of the resulting Rings infrastructure. RIP achieves this arrangement in an elegant and simple way. It arranges the nodes in the following sets: the network rings R, the network backbones B, the internal network nodes I , the network access points AP, and the mobile probes MP. A node can belong to one or more of these sets. Each member of R has a matching member in B, I , AP, and MP. A ring Ri consists of a backbone Bi and the associated internal nodes Ii. A backbone is a closed sequence of nodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002792_tte.2021.3081109-Figure25-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002792_tte.2021.3081109-Figure25-1.png",
        "caption": "Fig. 25. Prototype STPSPM machines. (a) Stator. (b) Rotor lacking 2 poles. (c) Rotor lacking 1 pole.",
        "texts": [
            " Authorized licensed use limited to: California State University Fresno. Downloaded on June 20,2021 at 01:06:31 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. A. Prototype SPM machines In this section, experiments are conducted to validate the numerical predictions. STPSPM machines are manufactured and tested, which are shown in Fig. 25. The parameters are the same with the 10-pole/12-slot STPSPM machines shown in Fig. 2(a). However, one of the rotors lacks two poles, which can be regarded as being symmetrically demagnetized, whilst the other one lacks one pole, which can be regarded as being unsymmetrically demagnetized. Firstly, the back-EMF waveforms for two STPSPM machines are measured. As shown in Fig. 26, the measured back-EMF generally matches well with the FE predictions. In addition, the cogging torque waveforms for two STPSPM machines are also measured, where the method in [26] is employed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003238_s12206-021-0829-0-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003238_s12206-021-0829-0-Figure3-1.png",
        "caption": "Fig. 3. Fan-driven gearbox lubrication mechanism.",
        "texts": [
            " The torque transmission bracket bears the gravity of the gearbox, and torque is generated when the gears mesh. To improve the oil-return capability of the fan-driven gearbox, the influence law of the planetary gear hole parameters on the heat transfer performance of the gearbox is determined by opening the planetary gear\u2019s undercut grooves and developing a design method for the planetary gear hole parameters, as shown in Fig. 2. The lubrication method for the fan-driven gearbox is illustrated in Fig. 3. One of the oil separators is equipped with eight nozzles. The torsion support, input shaft, and output shaft that exert minimal influences on heat production and heat transfer are omitted from the model due to the complex internal structure of the fan-driven gearbox. Considering that the fan-driven gearbox adopts a five-way shunt structure and is structurally symmetrical, a one-fifth gearbox is adopted for modeling in the CFD simulation. During meshing, a tetrahedral mesh that is highly adaptable to a complex model is used, and the meshes of the sun gear tooth surface, planetary gear tooth surface, ring gear tooth surface, and bearing surface are refined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure14.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure14.6-1.png",
        "caption": "Fig. 14.6 A free body diagram of the aerodynamic forces on the control actuation system of a missile. These terms and diagram are used for analysis of the forced response of the short-period approximation",
        "texts": [
            " Stability and frequency of the short period is dictated by the value of Cm\u03b1 . A negative value indicates a stable mode, and is confirmed by the center of gravity\u2019s forward position with respect to the vehicle\u2019s center or pressure The mode\u2019s damping is reliant on the magnitude of Cmq , and the vehicle\u2019s acceleration effectiveness is determined by Cz\u03b1 . All of the terms required for free response analysis are conceptualized in Fig. 14.5. The forced response of the system (u = 0) requires the stability derivative terms from Fig. 14.5 and the control derivatives from Fig. 14.6. It is important to think of the fins as a moment generators and not force generators. The control derivative term Cm\u03b4 is responsible for producing that moment. It is a product of the force generated at the fin and the moment arm from fin center of pressure to the vehicle\u2019s center of gravity. Provided there is a fixed moment arm length, increasing the surface areas of the actuation fins will lead to greater rotational authority. However, this comes at a cost. Increasing the force also increases the Cz\u03b4 term"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002504_j.ijmecsci.2021.106392-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002504_j.ijmecsci.2021.106392-Figure9-1.png",
        "caption": "Fig. 9. Forward upright motion of the torus.",
        "texts": [],
        "surrounding_texts": [
            "I e of the torus, and contains the hoop of equivalent radius. The contact point o ntre of the torus tube, and \ud835\udc3c is the lowest point of the equivalent hoop.\n\u2032\u27e9, \u27e8\ud835\udc4b \u2032\u2032\ud835\udc4c \u2032\u2032\ud835\udc4d \u2032\u2032\u27e9 and the body frame 2 \u27e8\ud835\udc4b 2 \ud835\udc4c 2 \ud835\udc4d 2 \u27e9 shown in Fig. 8 is given by\nE an be computed as:\n\ud835\udc93 (91)\nw\n\ud835\udc93 (92)\nwithout slipping. Thus, the velocity of the contact point can be computed\nf\n\ud835\udc97 (93)\nN nholonomic constraints are:\n\ud835\udc6a (94)\nG e system is \ud835\udc5b \ud835\udc54 = 3 . The equations of motion of the torus constitute a nonlinear i the nonholonomic constraints, results in the following index-1 DAE system: ( (95)\nw\n) ) sin 2 ( \ud835\udf19)\n\ud835\udc6b\n\ud835\udc78\nos ( \ud835\udf13 ) cos ( \ud835\udf19) \u0307 sin ( \ud835\udf13 ) cos ( \ud835\udf19) ?\u0307? sin ( 2 \ud835\udf19) n ( \ud835\udf19) cos ( \ud835\udf19) \u239e \u239f \u239f \u239f \u239f \u239f \u23a0 ,\n\ud835\udc85 (96)\nI r and a vertical axis perpendicular to the torus centre, respectively:\n\ud835\udc3c (97)\nn this way, the plane \ud835\udf0b\ud835\udc5a shown in Fig. 8 corresponds to the middle plan f the torus with the ground is denoted by \ud835\udc36; point \ud835\udc43 represents the ce\nAs in the hoop, the orientation of the intermediate frames \u27e8\ud835\udc4b \u2032\ud835\udc4c \u2032\ud835\udc4d\nqs. (70) . The absolute position vectors of \ud835\udc36 and the centre of mass \ud835\udc3a c\n\ud835\udc36 = ( \ud835\udc65 \ud835\udc36 \ud835\udc66 \ud835\udc36 0 )T , \ud835\udc93 \ud835\udc3a = \ud835\udc93 \ud835\udc36 + \ud835\udc93 \ud835\udc36\ud835\udc43 + \ud835\udc93 \ud835\udc43\ud835\udc3a ,\nith\n\ud835\udc36\ud835\udc43 =\n( 0 0 \ud835\udc4e )T , \ud835\udc93 \ud835\udc43\ud835\udc3a = \ud835\udc79 \u2032\u2032(0 0 \ud835\udc4f )T .\nThe nonholonomic constraints arise from the assumption of rolling\nrom Eq. (72) , obtaining:\n\ud835\udc36 = \u239b \u239c \u239c \u239c \u239d \ud835\udc63 \ud835\udc36 \ud835\udc65 \ud835\udc63 \ud835\udc36 \ud835\udc66 \ud835\udc63 \ud835\udc36 \ud835\udc67 \u239e \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239d ?\u0307? \ud835\udc36 \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) \u0307\ud835\udf03 cos ( \ud835\udf13 ) \u2212 \ud835\udc4e ?\u0307? sin ( \ud835\udf13 ) ?\u0307? \ud835\udc36 \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) \u0307\ud835\udf03 sin ( \ud835\udf13 ) + \ud835\udc4e ?\u0307? cos ( \ud835\udf13 ) 0 \u239e \u239f \u239f \u23a0 . ote that the \ud835\udc4d-component of the velocity is zero and, therefore, the no\n\ud835\udc5b\u210e ( \ud835\udc99 , ?\u0307? ) =\n(\n\ud835\udc63 \ud835\udc36 \ud835\udc65 \ud835\udc63 \ud835\udc36 \ud835\udc66\n)\n=\n(\n0 0\n)\n.\niven that \ud835\udc5b = 5 , \ud835\udc59 = 2 and \ud835\udc5a = 0 , the number of degrees of freedom of th ndex-2 DAE system which, by differentiating once with respect to time\n\ud835\udc74 ( \ud835\udc99 ) \ud835\udc6b T ( \ud835\udc99 ) \ud835\udc6b ( \ud835\udc99 ) \ud835\udfce\n) (\n?\u0308? \ud835\udf40\n)\n=\n(\n\ud835\udc78 \ud835\udc54 ( \ud835\udc99 ) + \ud835\udc78 \ud835\udc63 ( \ud835\udc99 , ?\u0307? ) \u2212 \ud835\udc85 ( \ud835\udc99 , ?\u0307? )\n)\n,\nith\n\ud835\udc74 ( \ud835\udc99 ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239d\n\ud835\udc5a 0 \ud835\udc5a\ud835\udc4f cos ( \ud835\udf13) sin ( \ud835\udf19 0 \ud835\udc5a \ud835\udc5a\ud835\udc4f sin ( \ud835\udf13) sin ( \ud835\udf19\n\ud835\udc5a\ud835\udc4f cos ( \ud835\udf13) sin ( \ud835\udf19) \ud835\udc5a\ud835\udc4f sin ( \ud835\udf13) sin ( \ud835\udf19) \ud835\udc3c \ud835\udc51 + ( \ud835\udc3c \ud835\udc5d \u2212 \ud835\udc3c \ud835\udc51 + \ud835\udc5a\ud835\udc4f 2 ) \ud835\udc5a\ud835\udc4f sin ( \ud835\udf13) cos ( \ud835\udf19) \u2212 \ud835\udc5a\ud835\udc4f cos ( \ud835\udf13) cos ( \ud835\udf19) 0\n0 0 \ud835\udc3c \ud835\udc5d sin ( \ud835\udf19) \ud835\udc5a\ud835\udc4f sin ( \ud835\udf13) cos ( \ud835\udf19) 0 \u2212 \ud835\udc5a\ud835\udc4f cos ( \ud835\udf13) cos ( \ud835\udf19) 0\n0 \ud835\udc3c \ud835\udc5d sin ( \ud835\udf19) \ud835\udc3c \ud835\udc51 + \ud835\udc5a\ud835\udc4f 2 0\n0 \ud835\udc3c \ud835\udc5d \u239e \u239f \u239f \u239f \u239f \u239f \u23a0 ,\n( \ud835\udc99 ) =\n(\n1 0 0 \u2212 \ud835\udc4e sin ( \ud835\udf13 ) \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) cos ( \ud835\udf13 ) 0 1 0 \ud835\udc4e cos ( \ud835\udf13 ) \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) sin ( \ud835\udf13 )\n)\n,\n\ud835\udc54 ( \ud835\udc99 ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239d\n0 0 0\n\ud835\udc5a\ud835\udc54\ud835\udc4f sin ( \ud835\udf19) 0 \u239e \u239f \u239f \u239f \u239f \u239f \u23a0 , \ud835\udc78 \ud835\udc63 ( \ud835\udc99 , ?\u0307? ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239d\n\ud835\udc5a\ud835\udc4f ( ?\u0307? 2 + ?\u0307?2 ) sin ( \ud835\udf13 ) sin ( \ud835\udf19) \u2212 2 \ud835\udc5a\ud835\udc4f \u0307\ud835\udf13 ?\u0307? c \u2212 \ud835\udc5a\ud835\udc4f ( ?\u0307? 2 + ?\u0307?2 ) cos ( \ud835\udf13 ) sin ( \ud835\udf19) \u2212 2 \ud835\udc5a\ud835\udc4f \u0307\ud835\udf13 \ud835\udf19\n\u2212 \ud835\udc3c \ud835\udc5d ?\u0307??\u0307? cos ( \ud835\udf19) \u2212 ( \ud835\udc3c \ud835\udc5d \u2212 \ud835\udc3c \ud835\udc51 + \ud835\udc5a\ud835\udc4f 2 ) ?\u0307?\n\ud835\udc3c \ud835\udc5d ?\u0307? ?\u0307? cos ( \ud835\udf19) + ( \ud835\udc3c \ud835\udc5d \u2212 \ud835\udc3c \ud835\udc51 + \ud835\udc5a\ud835\udc4f 2 ) ?\u0307?\n2 si \u2212 \ud835\udc3c \ud835\udc5d ?\u0307? ?\u0307? cos ( \ud835\udf19)\n( \ud835\udc99 , ?\u0307? ) =\n(\n\u2212 \u0307\ud835\udf13 ( \ud835\udc4e ?\u0307? cos ( \ud835\udf13 ) \u2212 ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) \u0307\ud835\udf03 sin ( \ud835\udf13 ) ) + \ud835\udc4e ?\u0307??\u0307? cos ( \ud835\udf13 ) sin ( \ud835\udf19) \u2212 \u0307\ud835\udf13 ( \ud835\udc4e ?\u0307? sin ( \ud835\udf13 ) + ( \ud835\udc4f + \ud835\udc4e cos ( \ud835\udf19) ) \u0307\ud835\udf03 cos ( \ud835\udf13 ) ) + \ud835\udc4e ?\u0307??\u0307? sin ( \ud835\udf13 ) sin ( \ud835\udf19) ) .\nn Eq. (96) , \ud835\udc3c \ud835\udc51 and \ud835\udc3c \ud835\udc5d represent the moments of inertia about a diamete\n= 1 \ud835\udc5a ( 5 \ud835\udc4e 2 + 4 \ud835\udc4f 2 ) , \ud835\udc3c = 1 \ud835\udc5a ( 3 \ud835\udc4e 2 + 4 \ud835\udc4f 2 ) .\n\ud835\udc51 8 \ud835\udc5d 4",
            "3\nexpressed as:\n\u23a7\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23a9 (98)\nw ference motion is equivalent to that of the hoop described by equations (76) , u q. (98) is given by \ud835\udc99 0 ( \ud835\udc61 ) = \ud835\udec00 \ud835\udc61, with \ud835\udec00 being the constant vector:\n\ud835\udec0 (99)\nT tion can be computed with Eq. (8) , leading to:\n\ud835\udf40 (100)\nht forward motion. Using the same coordinate partition as in the hoop case,\nt\n\ud835\udc99 (101)\nn can be expressed as a fist order system of the form \u0307\u0303\ud835\udc7f = \ud835\udc71 ?\u0303? . Developing\nE\n\ud835\udc71\n0 0 0 0 0 0 1 0 0\n) 0 0 0\n0 0 0 0 0 0 \ud835\udc4e + \ud835\udc4f 0 0 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 . (102)\neters:\n\ud835\udf0f (103)\nw p in Eq. (81) , the correponding set of eigenvalues is:\n\ud835\udf40 (104)\nw and \ud835\udf01 ( \ud835\udf02, \ud835\udf08) are given by:\n\ud835\udf11\n(105)\ncannot be further reduced using the procedure described in Subsection 2.4 ,\ns\n.3.2. Linearization in ODE form\nThe forward upright motion with constant speed of the torus can be\n\ud835\udc65 0 \ud835\udc36 ( \ud835\udc61 ) = \u03a9( \ud835\udc4e + \ud835\udc4f ) \ud835\udc61, \ud835\udc66 0 \ud835\udc36 ( \ud835\udc61 ) = 0 , \ud835\udf13 0 ( \ud835\udc61 ) = 0 , \ud835\udf190 ( \ud835\udc61 ) = 0 , \ud835\udf030 ( \ud835\udc61 ) = \u03a9\ud835\udc61,\nhere \u03a9 is, in this case, the rotational speed of the torus. Note that this re sing the relation of Eq. (90) . In vector form, the reference motion of E\n0 = ( \u03a9( \ud835\udc4e + \ud835\udc4f ) , 0 , 0 , 0 , \u03a9 )T .\nhe values of the Lagrange multipliers corresponding to this refence mo\n0 = ( \ud835\udf060 1 \ud835\udf060 2 )T = ( 0 0 )T .\nFigure ( 9 ) represents an arbitrary position of the torus during uprig\nhe sets of independent and dependent coordinates are:\n\u0303 \ud835\udc4e\ud835\udc56 =\n( ?\u0303? ?\u0303? \ud835\udf03 )T , ?\u0303? \ud835\udc51\ud835\udc51 = ( ?\u0303? \ud835\udc36 ?\u0303? \ud835\udc36 )T .\nDefining ?\u0303? = ( ?\u0303? \ud835\udc4e\ud835\udc56 \u0307\u0303\ud835\udc99 \ud835\udc4e\ud835\udc56 ?\u0303? \ud835\udc51\ud835\udc51 )T , the linearized equations of motio\nq. (32) , the resulting Jacobian matrix is given by:\n= \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d\n0 0 0 1 0 0 0 0 0 1 0 0 0 0 0\n0 \ud835\udc50\ud835\udc5a 1 . 15 \ud835\udc52\ud835\udc5a 0 0 0 0 \u2212\n2\u03a9 ( 3 \ud835\udc4e 2 +4 \ud835\udc4f 2\n5 \ud835\udc4e 2 +4 \ud835\udc4f 2\n0 8 \ud835\udc54\ud835\udc4f 13 \ud835\udc4e 2 +16 \ud835\udc4e\ud835\udc4f +12 \ud835\udc4f 2 0 2\u03a9\n( 7 \ud835\udc4e 2 +8 \ud835\udc4e\ud835\udc4f +8 \ud835\udc4f 2 ) 13 \ud835\udc4e 2 +16 \ud835\udc4e\ud835\udc4f +12 \ud835\udc4f 2 0\n0 0 0 0 0 0 \ud835\udc50\ud835\udc5a 1 . 15 \ud835\udc52\ud835\udc5a 0 0 0 0 0 \u03a9( \ud835\udc4e + \ud835\udc4f ) 0 0 0 \u2212 \ud835\udc4e\nThe size of \ud835\udc71 is (2 \ud835\udc5b \u2212 \ud835\udc5a \u2212 \ud835\udc59) \u00d7 (2 \ud835\udc5b \u2212 \ud835\udc5a \u2212 \ud835\udc59) = 8 \u00d7 8 . Introducing the param\n=\n\u221a\n\ud835\udc4e + \ud835\udc4f\n\ud835\udc54 , \ud835\udf08 = \u03a92 ( \ud835\udc4e + \ud835\udc4f ) \ud835\udc54 ,\nhich, by virtue of Eq. (90) , are equivalent to those defined for the hoo\nS = ( \ud835\udfce 1\u00d76 \ud835\udf061 S \ud835\udf062 S )T ,\nith \ud835\udf061 S = \ud835\udf11 ( \ud835\udf02, \ud835\udf0f) \u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08) and \ud835\udf062 S = \ud835\udf11 ( \ud835\udf02, \ud835\udf0f) \u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08) . The functions \ud835\udf11 ( \ud835\udf02, \ud835\udf0f)\n( \ud835\udf02, \ud835\udf0f) =\n2\n\ud835\udf0f \u221a ( 5 \ud835\udf022 + 4 )( 13 \ud835\udf022 + 16 \ud835\udf02 + 12 ) , \ud835\udf01 ( \ud835\udf02, \ud835\udf08) = ( 10 \ud835\udf022 + 8 ) ( 1 + \ud835\udf02) \u2212 \ud835\udf08 ( 21 \ud835\udf024 + 24 \ud835\udf023 + 52 \ud835\udf022 + 32 \ud835\udf02 + 32 ) .\nAs in the hoop case, the linearized equations of motion of the torus\nince there are no holonomic constraints.",
            "Table 3 Comparison of the results using the linearization approaches with the rolling torus.\nRolling Torus \ud835\udc5b = 5 , \ud835\udc5a = 0 , \ud835\udc59 = 2\nProcedure 1 Lin . DAE form\nProcedure 2 Lin . ODE form\nNull eigenvalues due to the procedure\n2 ( \ud835\udc5a + \ud835\udc59 ) = 4 \ud835\udc5a + \ud835\udc59 = 2\nNull eigenvalues due to the example\n4 4\nNonzero eigenvalues \ud835\udf061 S = \ud835\udf11 ( \ud835\udf02, \ud835\udf0f)\n\u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08)\n\ud835\udf062 S = \u2212 \ud835\udf11 ( \ud835\udf02, \ud835\udf0f) \u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08)\n\ud835\udf061 S = \ud835\udf11 ( \ud835\udf02, \ud835\udf0f) \u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08)\n\ud835\udf062 S = \u2212 \ud835\udf11 ( \ud835\udf02, \ud835\udf0f) \u221a \ud835\udf01 ( \ud835\udf02, \ud835\udf08)\nTotal number of eigenvalues\n2 \ud835\udc5b = 10 2 \ud835\udc5b \u2212 \ud835\udc5a \u2212 \ud835\udc59 = 8\nEfficiency Excellent Excellent\n3\ne linearization procedures with the rolling torus.\nth the approach employed is reduced from 2 ( \ud835\udc5a + \ud835\udc59 ) = 4 with the Procedure\n1 2 with the linearization in ODE form (Procedure 2), due to the presence of t s remain unchanged irrespective of the procedure used, existing four null e orrespond to the forward motion as a rigid body, whereas the remaining two n coupled, second-order, linear system of differential equations, involving the y\n(106)\ntained using Procedure 2 for simplicity, are given by:\n\ud835\udf53\n\ud835\udf53\n(107)\nw\n8 \ud835\udc4f 5 ) ,\n12 \ud835\udc4f 2 ) .\n(108)\nT ngles.\nnerates to the hoop, the nonzero eigenvalues \ud835\udf061 S , \ud835\udf06 2 S of the torus shown in\nE n-null eigenvalues, the boundary that determines the linear stability of the f\n\ud835\udf01 (109)\nw stabilization occurs. Using Eqs. (105) and (109) , one obtains:\n\ud835\udf08 (110)\nD copic stabilization is achieved by \u03a9\ud835\udf02\ud835\udc50 , and using Eqs. (85) , (90), (103) and ( op \u03a9\ud835\udc50 is found:\n\u03a9 (111)\nw\n\ud835\udf12 (112)\nF function, with \ud835\udf12( 0 ) = 1 . Therefore, it can be seen that the toroidal geometry r rotational speed of the torus verifies \u03a9\ud835\udf02\ud835\udc50 < \u03a9\ud835\udc50 \u2200\ud835\udf02 \u2260 0 . aracteristic time of \ud835\udf0f = 1 s . The critical rotational speed of the hoop with e \u2215s . As in the hoop, the torus presents a hyperbolic equilibrium for \u03a9 < \u03a9\ud835\udf02\ud835\udc50 , w = \u03a9\ud835\udf02\ud835\udc50 , the eigenvalues coalesce, so that \ud835\udf061 S = \ud835\udf062 S = 0 , and the system remains u and the real eigenvalues turn into a complex conjugate pair, leading to an e d complex parts of the eigenvalues \ud835\udf061 S , \ud835\udf06 2 S with the rotational speed \u03a9. The c s also highlighted.\n.3.3. Discussion and validation of the obtained results\nTable 3 shows a detailed breakdown of the results obtained using th As in the hoop case, the number of null eigenvalues associated wi\n(due to the existence of \ud835\udc5a + \ud835\udc59 = 2 dependent coordinates) to \ud835\udc5a + \ud835\udc59 =\nwo nonholonomic constraints. Furthermore, the remaining eigenvalue igenvalues and two nonzero eigenvalues. Two of the null eigenvalues c ull and two nonzero eigenvalues are associated with the following two aw and lean angles of the torus:\n\u0308\u0303\ud835\udf13 = \u2212\n2\u03a9 ( 3 \ud835\udc4e 2 + 4 \ud835\udc4f 2 ) 5 \ud835\udc4e 2 + 4 \ud835\udc4f 2 \u0307\u0303\ud835\udf19,\n\u0308\u0303\ud835\udf19 = 8 \ud835\udc54\ud835\udc4f 13 \ud835\udc4e 2 + 16 \ud835\udc4e\ud835\udc4f + 12 \ud835\udc4f 2 ?\u0303? +\n2\u03a9 ( 7 \ud835\udc4e 2 + 8 \ud835\udc4e\ud835\udc4f + 8 \ud835\udc4f 2 ) 13 \ud835\udc4e 2 + 16 \ud835\udc4e\ud835\udc4f + 12 \ud835\udc4f 2 \u0307\u0303\ud835\udf13.\nThe eigenmodes corresponding to the nonzero eigenvalues \ud835\udf061 S , \ud835\udf06 2 S , ob\n1 S = ( \ud835\udf13 0 \ud835\udf190 \ud835\udf030 ?\u0307? 0 ?\u0307?0 ?\u0307?0 \ud835\udc65 \ud835\udc36 0 \ud835\udc66 \ud835\udc36 0 )T ,\n= ( \u2212\n2 \ud835\udf05 \ud835\udefe\n\u221a\n\ud835\udefd \ud835\udf1a \u22122 \ud835\udefd \ud835\udefe 0 2\u03a9\ud835\udc3c \ud835\udc5d \ud835\udefd \ud835\udc3c \ud835\udc51 \ud835\udefe 4 \ud835\udefd \ud835\udefe\n\u221a\n\ud835\udefd \ud835\udf1a 0 0 1\n)T ,\n2 S = ( 2 \ud835\udf05 \ud835\udefe \u221a \ud835\udefd \ud835\udf1a \u22122 \ud835\udefd \ud835\udefe 0 2\u03a9\ud835\udc3c \ud835\udc5d \ud835\udefd \ud835\udc3c \ud835\udc51 \ud835\udefe \u22124 \ud835\udefd \ud835\udefe \u221a \ud835\udefd \ud835\udf1a 0 0 1 )T ,\nith \ud835\udefd, \ud835\udefe, \ud835\udf05 and \ud835\udf1a being the following constants: \ud835\udefd = \ud835\udc54 ( 10 \ud835\udc4e 2 \ud835\udc4f + 8 \ud835\udc4f 3 ) \u2212 \u03a92 (21 \ud835\udc4e 4 + 24 \ud835\udc4e 3 \ud835\udc4f + 52 \ud835\udc4e 2 \ud835\udc4f 2 + 32 \ud835\udc4e\ud835\udc4f 3 + 32 \ud835\udc4f 4 ) ,\n\ud835\udefe = \ud835\udc54 ( 20 \ud835\udc4e 3 \ud835\udc4f + 16 \ud835\udc4e\ud835\udc4f 3 ) \u2212 \u03a92 (3 \ud835\udc4e 5 \u2212 39 \ud835\udc4e 4 \ud835\udc4f \u2212 32 \ud835\udc4e 3 \ud835\udc4f 2 \u2212 88 \ud835\udc4e 2 \ud835\udc4f 3 \u2212 48 \ud835\udc4e\ud835\udc4f 4 \u2212 4\n\ud835\udf05 = \u03a9 ( 3 \ud835\udc4e 2 + 4 \ud835\udc4f 2 )( 13 \ud835\udc4e 2 + 16 \ud835\udc4e\ud835\udc4f + 12 \ud835\udc4f 2 ) , \ud835\udf1a = ( 5 \ud835\udc4e 2 + 4 \ud835\udc4f 2 )( 13 \ud835\udc4e 2 + 16 \ud835\udc4e\ud835\udc4f +\nhese eigenmodes correspond to a motion involving the yaw and lean a\nIt can be seen that, for the limit case \ud835\udf02 \u2192 0 , where the torus dege q. (104) become those of the hoop in Eq. (82) . Considering these no\norward upright motion with constant velocity of the torus is given by: ( \ud835\udf02, \ud835\udf08\u2217 ) = 0 ,\nhere \ud835\udf08\u2217 is the value of the nondimensional parameter \ud835\udf08 for which the\n\u2217 =\n( 10 \ud835\udf022 + 8 ) ( 1 + \ud835\udf02)\n21 \ud835\udf024 + 24 \ud835\udf023 + 52 \ud835\udf022 + 32 \ud835\udf02 + 32 .\nenoting the critical rotational speed of the torus for which the gyros\n110) , the following relation with the critical rotational speed of the ho\n2 \ud835\udf02\ud835\udc50 = \ud835\udf12( \ud835\udf02) \u03a92 \ud835\udc50 ,\nhere \ud835\udf12( \ud835\udf02) is a nondimensional function of the aspect ratio \ud835\udf02, given by:\n( \ud835\udf02) =\n4 ( 10 \ud835\udf022 + 8 ) ( 1 + \ud835\udf02)\n21 \ud835\udf024 + 24 \ud835\udf023 + 52 \ud835\udf022 + 32 \ud835\udf02 + 32 .\nig. 10 represents the function \ud835\udf12( \ud835\udf02) , being a monotonically decreasing esults in a stabilizing effect with respect to the hoop, since the critical\nFig. 11 (a) shows the root locus of a torus with \ud835\udf02 = 0 . 75 and a ch\nquivalent radius is \u03a9\ud835\udc50 = 0 . 5 rad\u2215s , and \u03a9\ud835\udf02\ud835\udc50 = \u221a \ud835\udf12( 0 . 75 ) \u03a9\ud835\udc50 \u2243 0 . 4834 rad ith two nonzero real eigenvalues of opposite sign, being unstable. For \u03a9 nstable. Lastly, when \u03a9 > \u03a9\ud835\udf02\ud835\udc50 , the gyroscopic stabilization is achieved lliptic equilibrium. Fig. 11 (b) represents the evolution of the real an ritical rotational speed \u03a9\ud835\udf02\ud835\udc50 , separating the unstable and stable zones, i"
        ]
    },
    {
        "image_filename": "designv11_63_0000124_012164-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000124_012164-Figure4-1.png",
        "caption": "Figure 4. Von mises stress",
        "texts": [],
        "surrounding_texts": [
            "Tegangan ekivalen yang digunakan metode Von-Mises. Berikut ini ilustrasi hasil analisis equivalent stressTegangan ekivalen maksimum terjadi di bagian las rangka bagian depan sebesar 16.32 MPa, kemudian tegangan ekivalen minimum sebesar 0.01 MPa."
        ]
    },
    {
        "image_filename": "designv11_63_0002914_s10846-021-01366-6-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002914_s10846-021-01366-6-Figure1-1.png",
        "caption": "Fig. 1 Frames",
        "texts": [
            " In general, a single rigid body represents the model of aerospace vehicles like fixed-wing UAVs [3] and quadrotors [29]. However, the presence of rotor and fuselage interaction in a helicopter restricts such assumption to model helicopter dynamics using single rigid body. Thus, it is important to obtain a proper mathematical model that can represent the rotor and fuselage interaction for a small scale helicopter. This paper considers a small scale fly-bar less helicopter having a fuselage and a main rotor (refer to Fig. 1). As we observe from Fig. 1, the fuselage is represented as a rigid body and a disk is used to model the helicopter rotor. A set of differential equations, each having first order, represents the motion model of main rotor flapping. We consider coupling between fuselage and rotor to model the presence of significant aerodynamic damping in the helicopter dynamics. As we consider a rigid body to represent fuselage subsystem, the configuration space containing all possible rotations, R, of the fuselage is denoted by the Special Orthogonal group SO(3), which is defined as SO (3) { R \u2208 R 3\u00d73| RRT = I3, det (R) = 1 } (1) where I3 denotes the third order identity matrix",
            " The group SO (3) is a Lie group having an algebraic group structure and matrix multiplication as the group operation. For more details on SO (3), please see [6]. The advantage to parameterize the attitude of a rigid body by using rotation matrices is that it is global and unique [8]. The motion equation to represent rotation of the rigid body fuselage is given as R\u0307 = R\u03c9\u0302 \u03c9\u0307 = \u2212J\u22121\u03c9\u0302J\u03c9 + J\u22121M (2) where R \u2208 SO (3) is the rotation matrix that transforms any vector from the body fixed frame (OB, xB, yB, zB) to the inertial reference frame (OG, xG, yG, zG) (refer to Fig. 1). The hat map \u02c6(.) : R 3 \u2192 so (3), where so (3) is the set of skew-symmetric matrices of order 3. We assume that fuselage inertia J \u2208 R 3\u00d73, expressed in the body frame, is constant and it is referred to its centre of mass. The angular velocity of fuselage is denoted as w = [p, q, r] \u2208 R 3, where p, q, and r are the components along the axes xB , yB , and zB , respectively. The main and tail rotor of the helicopter create external moment, M , on the fuselage. In the body fixed frame, this net moment acting on the fuselage is expressed as M = [l, m, n] \u2208 R 3, where l, m, and n are the roll, pitch, and yaw moments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001299_s00202-020-01062-y-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001299_s00202-020-01062-y-Figure11-1.png",
        "caption": "Fig. 11 Instrumented test bench 1: Motor to test 2: Brake 3: Tank 4: Recirculation pump 5: Valve 6: Flowmeter 7: Manometer 8: Temperature probe",
        "texts": [
            " The lifetime curve of the tested insulation material is illustrated in Fig.\u00a09. To validate the developed electrical and thermal models, a canned motor was designed and instrumented with temperature sensors in the stator coil heads, in the slots, on the motor housing and at the rear bearing as shown in Fig.\u00a010. An instrumented test bench has also been designed. It is composed of the motor to be tested, a tank containing the fluid that cools the motor and lubricates the bearings, temperature sensors, manometers, a flowmeter, a valve, and recirculation pump as shown in Fig.\u00a011. (34)(X) = 1 T , (35)B = E 2.303Rgas . In order to validate the model, several tests were conducted on an instrumented canned motor that allows identifying the mechanical losses and all the electrical parameters for the motor performance analysis. The mechanical losses were determined by the variable voltage method at a given cooling flowrate. A 5.5 KW motor connected in delta 380\u00a0V with different stator cans thicknesses and can material was tested. Figure\u00a011 shows the instrumented test bench. The calculated torque and efficiency given by Figs.\u00a012 and 13 show that the model results are in a good agreement with the measurements done with the test bench. As illustrated in Fig.\u00a014, the stator can losses are the most significant ones. PJs is for stator joule losses, Psc stator can losses, Prc rotor can losses, Pfs stator iron losses,Pfr rotor iron losses, Pjr rotor joule losses, and Pm mechanical losses. By comparing the measured end winding temperature rise above the fluid temperature with the end winding calculated temperature rise still above the fluid temperature for different operating points (different mechanical power and cooling flowrate, Table\u00a02), the experimental values are in line with the calculated ones"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002010_j.tws.2020.107334-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002010_j.tws.2020.107334-Figure6-1.png",
        "caption": "Fig. 6. Stress on the inner wall of the cylindrical shell at different positions.",
        "texts": [
            " The contact form of the actuator and the inner wall of the cylindrical shell is set as \u2018no separation\u2019; the contact form of the actuator and the circumferential sliding block is set as \u2018bonded\u2019. The static deformation simulation results can be obtained as shown in Fig. 5. In order to analyze the stress situation at the contact position between the actuator and the cylindrical shell more intuitively, we set the 12 o\u2019clock position in the circumferential direction of the cylindrical shell as 0\u25e6. The stress on the inner wall of the cylindrical shell at 30\u25e6, 45\u25e6 and 60\u25e6 is extracted respectively, as shown in Fig. 6. One can find from Fig. 6 that the peak value of the stress occurs at the contact position between the actuator and the inner wall of the cylindrical shell. It is easy to understand that the larger the peak value of the stress is, the more effectively the actuator is to the shell. Therefore, in order to investigate the influence of the thickness of the actuator on its actuation performance, firstly, the thickness of the middle of the actuator t is changed from 5 mm to 2 mm with intervals of 0.5 mm, while the axial width of the actuator b remains unchanged"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001197_s0263574720000533-Figure19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001197_s0263574720000533-Figure19-1.png",
        "caption": "Fig. 19. The illustration of envelope of the obstacles and the links.",
        "texts": [
            " However, computation of the B\u00e9zier curve will become expensive with the degree of the curve increasing. So we choose to elevate the B\u00e9zier curve of degree 4 to one of degree 5 as the backbone curve. The control points G0, G1, ....., G5 in Fig. 18(b) are computed according to {P0, P1, ...P4} using Eq. (47). Then G3 is moved to G\u2032 3 for adjusting the B\u00e9zier curve locally. Then, G0 = G\u2032 0, G1 = G\u2032 1, G2 = G\u2032 2, and G5 = G\u2032 5. G\u2032 4 is solved by Eqs. (11), (12), (15), and (16). For the steps of the avoiding obstacle, they can be expressed in Algorithm 1. These variables in Algorithm 1 are shown in Fig. 19. And the process of the avoiding obstacles is to loop the algorithm until all element of ds are not smaller than 0 or the loop time reaches the set maximum iteration time. \u03b7 in Algorithm 1 is a step size of the point movement. As described in Eq. (47), the flexibility of the backbone curve of the snake robot can be increased by adding more control points to the B\u00e9zier curve. When the backbone curve has more flexibility, the head of snake robot can avoid more obstacles in the environment at the same time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001915_icem49940.2020.9270831-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001915_icem49940.2020.9270831-Figure10-1.png",
        "caption": "Fig. 10: Torsional mode shape of the stator at 825 Hz",
        "texts": [
            " Nonetheless, the simulation shows good accordance with the measured eigenfrequencies, as indicated in table III. As recent investigations have shown, the first torsional mode shape [16] and the breathing mode shape [20] of the stator are of significance for the driver\u2019s cabin interior noise and need to be considered. As the torsional mode occurs only in the mounted state, the flange on which the motor is fixed needs to be taken into account. The elastic boundary condition allows the stator to twist against the flange. The mode shape calculated in such a way is shown in fig. 10. The asymmetric shape of the mode is due the non-equidistant distribution of the screw threads of the motor housing. Unfortunately, the mode shape is difficult to measure and depends on the flange. Starting with the spring constant given in [21], the boundary condition can be adjusted by editing this constant. Simply applying a fixed boundary 1188 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 16,2021 at 18:43:55 UTC from IEEE Xplore. Restrictions apply"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002578_j.jclepro.2021.126900-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002578_j.jclepro.2021.126900-Figure4-1.png",
        "caption": "Fig. 4. A) 3D render of the bladed sector gripping system; B) stator sector mounted on the gripping system.",
        "texts": [
            " To this end, a 3D model of the bladed sector was created using Solidworks\u00ae software, with which the gripping system was designed. The latter consists of a vertical shaft (fixed to the central pulley of the handling system so that the piece had only a rotary motion around the FB chamber axis), a Fig. 3. A) Schematisation of the FB system and B) experimental setup. The dimensions are ex inverter and control unit; (5) porous filter; (6) air supply. system for adjusting the position of the bladed sector, and a series of registers to allow the secure fixing of the component (see Fig. 4A). Once the gripping system was designed, the relative construction was entrusted to a mechanical workshop, which used an aluminium alloy so that the weight was limited and the components were easily workable, thus guaranteeing a low production cost. Fig. 4B shows the final gripping system. Surface roughness on aeroengine blades plays a critical role in the performance and, therefore, on the resulting efficiency of the blades (Priegue and Stoesser, 2017). It can derive from the manufacturing process and further increase during the operation. Moreover, the roughness on a particular blade also affects the wake flow and, therefore, the inflow conditions for the subsequent turbine stages (Mulleners et al., 2014). In this context, the technological feasibility of FB processing has been verified by analysing the final surfaces through the characterisation of their roughness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001802_j.snb.2020.129151-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001802_j.snb.2020.129151-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_63_0001158_j.ymssp.2020.107051-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure13-1.png",
        "caption": "Fig. 13. GSL lib",
        "texts": [
            " Similar variation in Y-direction is performed for different values of X thus producing a grid (Fig. 12(b)) over which the evaluations are performed. The contact forces at all teeth are thus evaluated at all the simulated positions of the rotor center. This procedure is repeated for different angles of orientation of rotor. This evaluation of forces requires considerable computational time, if performed sequentially, and thus these evaluations are performed in parallel using OpenMP, with a C++ implementation. At a given simulated position of the rotor (Fig. 13(a)), the radii of curvature of both the intersecting profiles are evaluated at the contact point. The penetration of profiles (Fig. 13(b)) is evaluated at different rollers using GSL libraries at all the contact points. At points of no contact, the corresponding gap height and equivalent gap length are evaluated and stored so that the variation of tooth tip gap features with micro-motion of rotor can be considered in the analysis. The gap between two TSVs acts as a passage for flow communication. The flow through this gap is considered as the case of flow through a circular pipe of equivalent gap length (Fig. 14). This equivalent gap length (lt) was evaluated as distance between points on the gears that satisfy h t \u00bc ht 1\u00fe \u00f0 \u00de where was found to be 0:36 based on both experiments and simulations individually [25]. It should be noted that the stator presented in Fig. 13(a) is a combination of stator body and rollers. To evaluate the contact forces more accurately, the contact between the rollers-stator body is also modelled as shown in Fig. 15(a). For a given roller, the net indentation between the stator and the rotor is evaluated (Fig. 13(b)), and this indentation, d, is distributed between the rotor-roller contact, d1, and the roller-stator body contact, d2, as shown in Fig. 15b (inset of a roller in Fig. 15 (a)) such that d \u00bc d1 \u00fe d2 and the forces (highlighted by arrows) evaluated at both the contacts are same so that the net force on the roller is nullified. A one-dimensional root finding algorithm [26] is employed to find the value of d1 and consequently, d2. Apart from the magnitude of the contact forces, both the direction and the arm-length of these forces with respect to the rotor center are also evaluated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000016_chicc.2019.8865860-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000016_chicc.2019.8865860-Figure2-1.png",
        "caption": "Fig. 2: Closed-Loop of One Leg.",
        "texts": [
            " To acquire the rotation matrix of the moving platform, the rotation Angle (\u03b1) around the Z-axis is called Roll, the rotation Angle (\u03b2) around the Y-axis is called Pitch, and the rotation Angle (\u03b3) around the X-axis is called Yaw. The rotation matrix of the moving platform relative to the base platform is as follows: RB p = Rz(\u03b3)Ry(\u03b2)Rx(\u03b1) (1) RB p = \u23a1 \u23a3c\u03b2c\u03b3 s\u03b1s\u03b2c\u03b3 \u2212 c\u03b1s\u03b3 c\u03b1s\u03b2c\u03b3 + s\u03b1s\u03b3 c\u03b2s\u03b3 s\u03b1s\u03b2c\u03b3 + c\u03b1s\u03b3 c\u03b1s\u03b2c\u03b3 \u2212 s\u03b1s\u03b3 \u2212s\u03b2 s\u03b1c\u03b2 c\u03b1c\u03b2 \u23a4 \u23a6 (2) To calculate the lengths of six legs, the closed-loop vector method shown in Fig. 2 can be used. Any coordinates of the moving platform can be transformed into a fixed coordinate system of the base platform by a rotation matrix. Therefore: qB i = RB p qp i (3) Further, the expression of the leg length vector can be ob- tained: Li = t+RB p qp i \u2212 bi (4) where i \u2208 [1, 6]. If the coordinates of the origin P of the moving system in the fixed system are (xp, yp, zp), the hinge coordinates of the base platform are (xbi, ybi, zbi), the hinge coordinates of the moving platform in the fixed system are (xpi, ypi, zpi), and the initial lengths of the six legs are l0i, the inverse kinematics solution can be further written as follows: li = fi(q) = fi(\u03b1, \u03b2, \u03b3, x, y, z) = [(xpi \u2212 xbi) 2 + (ypi \u2212 ybi) 2 + (zpi \u2212 zbi) 2]1/2 (5) where q represents the position and attitude of the platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001813_j.jterra.2020.09.002-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001813_j.jterra.2020.09.002-Figure1-1.png",
        "caption": "Fig. 1. Single wh",
        "texts": [
            " The experimental observation is used for an analysis of the wheel camber effect to the slope traversability. We utilize the particle image velocimetry (PIV) technique to visualize the soil flow and soil failure beneath the wheel. The rest of this paper is organized as follows: Section 2 introduces the experimental setup with the in-wheel camera. Section 3 describes the experimental result and discussions under various wheel camber angles and slip ratios. Section 4 summarizes the findings and future application of this study. The single wheel testbed (Fig. 1) has dimensions of 3500 mm in length, 600 mm in width, and 1200 mm in height. The testbed is Nomenclature a0; a1 Coefficients of the maximum normal stress angle (\u2013) b Wheel width (m) Fc Cornering force (N) Fl Lateral force on slope coordinate (N) Fx Drawbar pull on wheel coordinate (N) Fy Side force on wheel coordinate (N) Fz Normal force on wheel coordinate (N) r Wheel radius (m) s Slip ratio (\u2013) vx Wheel velocity (m/s) W Wheel load (N) hc Wheel camber angle (rad) hf Entry angle (rad) hm Maximum force angle (rad) hs Slope angle (rad) j Exit angle coefficient (\u2013) x Wheel angular velocity (rad/s) filled with the Toyoura sand that has an average particle diameter of 0",
            " Here, the wheel slip ratio is the state variable that expresses a magnitude of the wheel slip in the longitudinal direction defined as follows (Bekker, 1960): s \u00bc 1 vx rx \u00f01\u00de where vx is the translational velocity of the wheel, x is the wheel angular velocity, and r is the wheel radius. The slip ratio is realized by tuning the velocity of the traction motor to a given wheel angular velocity. The vertical wheel displacement is measured by the magnetic scale installed on the vertical slide-guide. A six-axis force/torque sensor on the above of the wheel measures the interaction forces generated at the wheel. The normal direction of the sensor always aligns to the normal direction of the wheel (Fig. 1(b)). The wheel camber angle mechanism is explained in Section 2.2. The inwheel camera is also introduced in Section 2.3. The wheel camber angle mechanism developed in this work is shown in Fig. 2. The motor controls the wheel camber angle via the worm gear/wheel mechanism. The worm gear mechanically prevents back-drive from the wheel forces so that the camber motor does not need to feed any current while it holds the desired camber angle. As noted in Section 2.1, the force sensor rotates as the camber angle rotates such that the wheel coordinate always matches with the coordinate of the force sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure8-1.png",
        "caption": "Fig. 8. (a) Force on rotor due to pressure inside one of the TSV, (b) Normal contact force and friction forces on rotor at one of the rollers.",
        "texts": [
            " The flow through the commutator lubricating gap and lateral lubricating gaps are evaluated by Gap module and is detailed in Section 3.4. The corresponding leakages obtained from the Gap module are included in the flow evaluation of fluid dynamics and force module. The pressure inside each TSV, evaluated using lumped parameter approach, is used to evaluate the corresponding force acting on the rotor F ! i as: F ! i \u00bc pTSV Sybx \u00fe Sxby \u00f06\u00de where Sx; Sy represent the projection of areas along X and Y axes (Fig. 8(a)), whose evaluation will be detailed in Section 3.2. This pressure force due to pressure inside one of the TSVs is shown in Fig. 8(a). Upon summation of these forces P F ! i , the net force that rotor experiences owing to the pressure inside all the TSVs can be obtained. Fig. 8(b) presents contact forces (normal and frictional) at one of the rollers. From the output of mechanical pre-processor module, which is detailed in section 3.3, normal contact force can be obtained. The normal contact force F ! c represents the normal force due to contact at one of the teeth. The frictional contact force, F ! f is evaluated as Ff \u00bc l Fc , where l is the friction coefficient, whose evaluation is discussed in the next sub-section. By evaluating all the forces instantaneously, the net force on the rotor can be evaluated as: F "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001158_j.ymssp.2020.107051-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001158_j.ymssp.2020.107051-Figure16-1.png",
        "caption": "Fig. 16. Exploded view of a portion of the orbit motor, highlighting the lubricating interfaces between Moving Elements (ME) and Stationary Elements (SE).",
        "texts": [
            " Upon evaluation of these quantities at all the possible contact points, the values are stored in a lookup table for subsequent use. Such table can be used to interpolate for contact force values, arm-lengths and normal angles at a given position of the rotor. This approach can be applied to any machine, but the values of Dxmax and Dxdiff are not known a priori and a choice of them can only be justified after obtaining the maximum deviation of the machine observed, after the implementation of fluid dynamics & force module. Fig. 16 represents an exploded view of the motor with the components numbered in accordance to Fig. 2. The gap module evaluates the leakages across the lubricating interfaces (arrows in Fig. 16) of the orbit motor. In particular, there are two moving elements (ME), the rotor and commutator, in conjunction with stationary elements (SE). Lubricating interfaces exist on either side of the orbiting rotor (4). Also, a lubricating interface exists between the commutator (6) and windows manifold (5). A sealing element is present between the commutator and the casing, thus effectively sealing the gap. The gap height of these lubricating interfaces is evaluated from the lengths of different components and by assuming equal gap height distribution on either side of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000729_humanoids43949.2019.9035004-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000729_humanoids43949.2019.9035004-Figure7-1.png",
        "caption": "Fig. 7. MEMS RPA assembly tool: A central circular Delrin part holds a hexagonal recess with protrusions for aligning the RPA housing to the center of the tool; this part rotates relative to the cam base on which the top spring manipulators ride to slowly pull back the springs.",
        "texts": [
            " With a beam width of 700 \u03bcm and a beam height of 300\u03bcm, each of the curved springs applies a force of roughly 0.1N when deflected 100 \u03bcm. Each spring is set 105 \u03bcm from center, to account for the foreshortening of the spring when it is actuated. Simulation of the spring using the commercial software COMSOL confirmed that at maximum deflection (200 \u03bcm, limited by a stop), the principal stresses remain below 150MPa (Fig. 6). We designed an assembly tool to facilitate the integration of the electrode grids to the housing (Fig. 7). In this fashion, there is no risk of added axial stress to the springs when the device is assembled, nor there is a risk of buckling the spring during disassembly. The springs have the same orientation such that all will expand or contract the same. A cam-like actuation was used to allow the spring tips to slowly be pulled back over a rotation of about 100\u00b0 to their limit stop. The MEMS RPA grids are then inserted one at a time using tweezers, and a rotation in the opposite direction closes the springs onto each electrode, latching them in perfect alignment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000591_978-3-030-36621-6_2-Figure13.19-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000591_978-3-030-36621-6_2-Figure13.19-1.png",
        "caption": "Fig. 13.19 Road coordinate system with corresponding velocities",
        "texts": [
            " If the acceleration of the preceding vehicle is small such that |ueq,i (t) + \u03b4\u0303i | \u2264 umax holds, then the formation can be maintained even if the vehicles are on different lanes. Similarly, a lane change can also be considered as a disturbance that arises from different curvatures. During a lane change, the lateral position of the vehicle on the road yi is changed in a smooth way due to a bounded lateral acceleration. Note that the states xi , vi in (13.56) are considered with respect to the longitudinal road coordinates as depicted in Fig. 13.19. For the longitudinal control in (13.56), it is assumed that the relative distance to other objects e\u0303x,i j = x j \u2212 xi and velocities vi , vi\u22121 are available with respect to the road coordinate system. The lateral velocity of the vehicle vy,i as shown in Fig. 13.19 is larger than zero during a lane change, and the velocity of the vehicle vcar,i is not equal to the road velocity vi in (13.56); specifically, vcar,i = \u221a v2i + v2y,i . (13.57) Assuming a point mass, the vehicle can be described in the x, y coordinate system by vi = x\u0307i = vcar,i cos(\u03b8i ) , vy,i = y\u0307i = vcar,i sin(\u03b8i ) , \u03b8\u0307i = \u03c9i , (13.58) where \u03b8i is the orientation with respect to the x-axis as depicted in Fig. 13.19 and \u03c9i is the angular velocity. The acceleration of the car v\u0307car,i = ucar,i and the acceleration ui in (13.56) differ as well. By differentiation of (13.58), the dynamics along the x-axis are v\u0307i = v\u0307car,i cos(\u03b8i ) \u2212 vcar,i sin(\u03b8i )\u03b8\u0307i = ucar,i cos(\u03b8i ) \u2212 y\u0307i \u03b8\u0307i = f (vi )ucar,i \u2212 di . (13.59) The orientation \u03b8i is assumed to be small, which is practicable on highways, thus cos(\u03b8i ) \u2248 1, sin(\u03b8i ) \u2248 0. Then, 0 < fmin \u2264 | f (vi )| \u2264 1.Assuming a bounded\u03c9i , the matched disturbance di is bounded, i",
            " Note that if the disturbance cannot be compensated, the position error to the preceding vehicle increases during the lane change due to a larger waylength, which is not safety-critical. Further analysis of theATH in presence of disturbances is subject to future work. In order to test the robustness of the algorithm with respect to lane changes or different curvatures, an example of a lane change is considered. As a sampling time, Ts = 0.001 s is used. Note that the path of the lane change in Fig. 13.20a is chosen to start rather abruptly on purpose in order to investigate the effects of different lateral accelerations. The velocity of the car as shown in Fig. 13.19 is larger during the lane change in order to maintain a constant velocity along the road as shown in Fig. 13.20b. The position errors of the two followers are presented in Fig. 13.20c, where no effects caused by the lane change can be detected. Note that the accelerations in Fig. 13.20e have been filtered with a first-order lag element with time constant \u03c4 = 0.01 s. The disturbances are bounded and small as shown in Fig. 13.20e. They can thus be rejected, and sliding can be maintained as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001138_978-3-030-48122-3-Figure6.9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001138_978-3-030-48122-3-Figure6.9-1.png",
        "caption": "Fig. 6.9 Schematic representation of a conformal cooling array employed in the inner refrigeration system belonged to an injection mold. Inner channels (2\u00a0mm in diameter) are manufactured through AM using DMLS",
        "texts": [
            " As light alloys in the manufacture of injection molds for the production of plastic parts, aluminum alloys are used, specifically those of aerospace grade. This is due to the fact that traditionally in those injection molds made from steel, the time required to cool a mold constitutes about 70% of the duty cycle for each part produced [3]. To further increase the productivity, the incorporation of conformed cooling channels is an alternative to improve the cooling time and the quality of the plastic part (Fig.\u00a06.9). Traditional subtractive manufacturing methods are used in the design and manufacture of cooling channels, while channels manufactured by means of AM provide more precise temperature control of the molding cavity during the injection cycle. The additive manufacturing and handling of metal powders allow the manufacture of parts of high interest in the tools industry [17]. This is because it allows considering architectures for new applications and new materials with interesting physical characteristics, which represent a competitive advantage in terms of saving machining time and increased performance in heat dissipation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001652_rcs.2180-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001652_rcs.2180-Figure7-1.png",
        "caption": "Fig. 7 The structure of the instrument used for suturing",
        "texts": [
            " The body coordinate system was established with the navel as the coordinate origin, the o-xy plane was parallel to the patient coronal plane and the z-axis was perpendicular to the paper and pointed outward. The wound position p1p2 was given by the general physiological structure herein, which can be obtained by edge extraction and 3D reconstruction of endoscope binocular images in practical surgery. Fig. 5 The simulation platform Fig.6 Sketch of the patient's physiological model in the duodenal repair operation A cc d A rt ic This article is protected by copyright. All rights reserved. The instrument used for suturing is shown in Fig. 7. The lengths of the instrument rod and the wrist were 300 mm and 20 mm, respectively, and the length of the end effectors was 15 mm. The needle used was O 1/2 8 x 20. As shown in Fig. 8, O represents that the cross section of the needle tip is circular, 1/2 represents that the body of the needle is curved in a semicircle, 8 represents that the diameter of the middle section of the needle body is 0.8 mm, and 20 represents that the distance between the two ends of the needle is 20 mm. Fig. 8 Diagram of the O 1/2 8 x 20 needle When suturing, the needle was rotated around its center to insert into the tissue, and the needle entry point and exit point were kept symmetrical to wound p1p2 [22-24] to shorten the path of the needle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure8-1.png",
        "caption": "Fig. 8. Temperature distribution for slot current density J =19.2 A(rms)/mm2",
        "texts": [],
        "surrounding_texts": [
            "The LPTN is used to provide a quick estimate of temperatures at different locations in the motor for different slot current densities. These were then compared against the temperatures obtained from CFD analysis to verify the validity of the LPTN model. Temperature rise in different portions of the machine using both CFD and lumped parameter network are shown in Table IV. It is visible that CFD results match closely with that of the LPTN model and the continuous output power for this 120 kW slotless machine using the proposed WELC method is 85 kW when the temperature rise is limited to 80 \u25e6C. This also corresponds to a current density of 23.3 A/mm2. This validates that the proposed WELC method is effective to further increase the maximum allowable continuous current density compared to the state-of-the-art cooling methods, which makes it possible to improve the power density as well. To analyze the effectiveness of thermal conductivity of the winding support material, the hotspot temperature of the winding is studied for different thermal conductivities ranging from 3 W/m-K (similar to Zirconia) to 18 W/m-K (similar to Silicon Nitride), as shown in Fig. 11. It is found that the continuous current density can be increased using materials with higher thermal conductivity. It is also found that, the benefits of higher thermal conductivity saturates at higher values."
        ]
    },
    {
        "image_filename": "designv11_63_0003208_s42835-021-00842-1-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003208_s42835-021-00842-1-Figure2-1.png",
        "caption": "Fig. 2 An exoskeleton interacting with the operator\u2019s lower limb during walking",
        "texts": [
            " In the past, these models could be obtained using standard rigid body formulations, but this solution is limited due to model errors, actuator dynamics, and un-modeled nonlinearities [32, 33]. Moreover, it is not feasible to apply the solution for control of the combined human-exoskeleton system since the system is especially subjected to unpredictable human\u2013robot interaction forces (torques). An alternative to analytically deriving the dynamic model as well as the interaction dynamics is to learn them. In swing phase, the human subject\u2019s leg including HUALEX is generally modeled as a 2-DOF serial link mechanism (Fig.\u00a02). The dynamics for this serial link are represented by the equation: where q \u2208 R2 is the generalized position vector comprising the hip and knee angles. The terms M(q) \u2208 R2\u00d72 , V(q) \u2208 R2 , (1) [ME(q) +MH(q)]q\u0308 + [VE(q, q\u0307) + VH(q, q\u0307)]q\u0307 + GE(q) + GH(q) + \ud835\udf0fF + \ud835\udf0fH = \ud835\udf0fA, G(q) \u2208 R2 represent the positive definite inertia matrix, Coriolis and centrifugal torque vector, and gravity torque vector, respectively, of the human subject (index H) and the exoskeleton (index E). The term \u2208 R2 of Eq.\u00a0(1) is the torques acting onto the exoskeleton",
            " The exoskeleton is controlled by a simple master\u2013slave PD controller with high gains in the first N1 iterations of the trajectory and then switched to the partitioned controller with the model being learnt. Half of the observed data was used for training the model and half for testing and estimating the accuracy of the learned model. In order to obtain this adequate movement data, the design of the position master\u2013slave controller for collecting data depends on several setup issues, for example, a proper mechanical connection between the master (human) and the slave (exoskeleton) as exampled in Fig.\u00a02. Moreover, only data in swing phase collected for learning process requires a suitably sensory setup for phase detection as discussed in the next section (Fig.\u00a04). The control scheme discussed in Sect.\u00a02 is motivated through simulations and experiments on HUALEX in this section. To deploy MLPC, the simplified configuration of HUALEX in swing phase is considered when the operator's leg with the exoskeleton is regarded as a two DOFs multi-link pendulum. From bio-mechanical studies of human behavior, HUALEX was designed as an anthropomorphic configuration [29, 44]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001770_dvm49764.2020.9243883-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001770_dvm49764.2020.9243883-Figure1-1.png",
        "caption": "Fig. 1. The hydraulic valve on the left was machined from steel, and then zinc plated for corrosion resistance. The redesigned valve (right) was produced by 3D printing in stainless steel.",
        "texts": [
            " Each material has differently heat and cool temperatures, depending on the speed and input energy as it interacts with the metal powder. The thermal state is inducing microstructure formation that provide corresponding material properties: strength, stiffness, and elasticity. Unfortunately, many SLM materials require post-processing heat treatment because of the stresses and hot cracking, which could occur during SLM process. Many types of hydraulic and pneumatic components already produced by additive technologies. For example, Aidro printed a stackable hydraulic valve with design improving (Fig.1) [1]. The 3D-printed valve weighs 60% less but maintains the same strength as the machined part (Courtesy of Aidro SrL). In order to have a lightweight hydraulic valve, the valve body has been redesigned to be 3D printed [2]. Authorized licensed use limited to: Rutgers University. Downloaded on May 18,2021 at 15:37:54 UTC from IEEE Xplore. Restrictions apply. 978-1-7281-7526-3/20/$31.00 \u00a92020 IEEE Additive manufacturing implementation can provides a rated flow improvement, less leakage, less volume, weight, and contains fewer parts, which simplifies fabrication of hydraulic and pneumatic components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000736_icpedc47771.2019.9036708-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000736_icpedc47771.2019.9036708-Figure2-1.png",
        "caption": "Fig. 2. Wind turbine operating regions",
        "texts": [
            " Therefore, there should be some kind of rigid controllers are required to extract max power and stiff voltage with frequency. The wind turbine of the WECS is coupled to the generator either through gearbox or direct drive. The available kinetic energy from the wind is converted into mechanical energy from through wind turbine. The generator converts mechanical energy from wind turbine shaft into electrical energy. The region of operation of wind turbine is classified into four regions as depicted in figure 2. The wind turbines are allowed to operate only between the cut-in speed and cut-out speed. Beyond cut-out speed the wind turbine should stopped against damage of turbine and generator. In region R1, the wind turbine should be stopped an isolated from grid to avoid turbine being rotated by generator. In the region R2, the wind turbine starts power at rated speed and also maximum power will be generated. Moreover, it should be regulated to extract maximum power from the available wind speed. II. WECS CONFIGURATION A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001088_b978-0-12-819972-5.00006-9-Figure6.6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001088_b978-0-12-819972-5.00006-9-Figure6.6-1.png",
        "caption": "Figure 6.6 Wi-Fi drone communication.",
        "texts": [
            " Trajectory planning is used to execute air take-off position, target position, goals, and tasks of the flight [21]. This enhances the flight movement of drones especially in environments with many obstacles \u2022 Automatic navigation system Simulating drones with AI greatly increases its speed to over 90 miles per hour, which makes the drones more accurate than an onboard pilot [22]. Simulating the automatic navigation system reduces the risk of poor navigation performance [15] and enables the system make use of precise units for the calculation of measurements (Fig. 6.6). \u2022 Communication System Drones are the capable of transmitting data such as speed, direction, fixed points [10] with the aid of Wi-Fi for as long, as they are within the same transmission radius [23], which actually reduces the unit of its memory and power. They communicate with other drones, the ground station, and the user applications involved. Parameters used for communication simulation includes bandwidth, latency, throughput, and data rate [10]. 4 Simulation environments The Table 6.1 describes some simulation environment that can be used test controllers and algorithms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002844_09544062211015784-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002844_09544062211015784-Figure1-1.png",
        "caption": "Figure 1. Exploded and details dimensions of strain wave gear. (a) Exploded view of harmonic drive: (i) Strain wave generating cam; (ii) circular spline; (iii) flex spline. (b) Geometrical representation of each component.",
        "texts": [
            " Asymmetric teeth, rotary actuator, strain wave gear, tooth profile, precision drive Date received: 23 November 2020; accepted: 16 April 2021 Effective power transmission and high load carrying capacity is usually an influential aspect in general gear design. This is always possible with compact size strain wave gear (SWG). The strain wave gear more commonly also known as harmonic drive (HD) is a precision drive that possess high load carrying capacity even at high transmission ratio. It is a noncircular gear (Figure 1) having vast applications in precisions instruments, robotics, avionics etc. as rotary actuator. It consists of internal toothed rigid circular spline (CS) which is stationery unit, external toothed flexible type flex spline (FS) cup which is output unit and oval shape strain wave generating (SWG) cam which is input unit. In general, for avoiding teeth interference, circular gear is preferred over non-circular and also number of teeth difference between mating gears is preferably being kept more than 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000834_crc.2019.00024-Figure14-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000834_crc.2019.00024-Figure14-1.png",
        "caption": "Fig. 14. Level sensor",
        "texts": [
            " connected to Faulhaber motion controllers (MCDC3003-S and MCDC3006-S), which were also connected to the laptop control the robot using the Faulhaber controllers. The wheelchair was equipped with the same accelerometer and ultrasonic sensor systems (Fig. 13 (b)). These sensor systems on the wheelchair were also connected to the laptop on the robot using a wireless Zigbee module. Two batteries and two motion controllers were added for the electric drive unit of the wheelchair, and connected to the robot\u2019s laptop via a Zigbee module. Level sensor systems were set on the robot and wheelchair (Fig. 14) to detect the inclination of each vehicle, and they were used to detect the arrival of the wheelchair or robot on level ground, that is, the end of the climbing operation (see Section III). In this study, stages 1 and 2 signify the processes in which the wheelchair front and back wheels, respectively, ascend the step. Similarly, stages 3 and 4 signify the processes in which the robot front and other wheels, respectively, climb a step. Both vehicles move automatically through the step ascent process illustrated in Fig",
            " <7> The robot continues to push the wheelchair so that the rear wheels of the wheelchair climb up onto the step. The 78 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 06:18:12 UTC from IEEE Xplore. Restrictions apply. robot supports the wheelchair during this process to prevent the wheelchair from tipping over backward. <8> Both vehicles move forward, and the chair rear wheels climb the step. The level sensor on the wheelchair detects the completion of the wheelchair climbing onto the step (Fig. 14). Stage 3 <9> The system retracts the front and rear wheel mechanisms of the robot (Fig. 5). The arms are spread apart (Figs. 11 (a) and (b)), the touch sensors on the forearm links detect the contact with the stopper of the wheelchair (Fig. 7), and the manipulator forearm links are inserted into the stopper (between the front and rear bars). The chair stops, the robot moves forward, and the manipulator forearm links come into contact with the front bars of the wheelchair stopper (Fig. 10). <10> The robot continues to push the wheelchair using the forearm links, and the front wheels of the robot are lifted until the accelerometer system on the robot detects the robot inclination indicating that both vehicles are in a position where the robot\u2019s front wheels are on the step",
            " <12> Both vehicles move forward, and the front wheels of the robot are placed on the upper level of the step. Stage 4 <13> Both vehicles continue to move forward. <14> The robot middle wheels come into contact with the step. The wheelchair pulls the robot. The robot rear wheels are lowered and help the middle wheels climb the step. <15> Both vehicles continue to move forward. The robot rear wheels remain deployed. <16> When the robot middle wheels have reached the upper level of the step, the level sensor on the robot detects the end of the step climbing of the robot (Fig. 14). Both vehicles stop and the robot rear wheels are retracted (folded upward) (Fig. 5). We experimented with the proposed system under an environment with a step 120 mm high and a friction coefficient \u03bc = 0.72 (the floor and step material were wood; Fig. 16). The wheelchair user was an able-bodied adult male (age 20, weight 50 kg), and the experiment was done on one floor of the National Institute of Technology, Toyama College. The velocities of both vehicles were constant (0.76 km/h). In stage 1, the wheelbase between the robot front and middle wheels, WB f (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001470_2050-7038.12588-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001470_2050-7038.12588-Figure10-1.png",
        "caption": "FIGURE 10 Cogging torque variation",
        "texts": [
            " The magnet's mechanical angle of rotation can be calculated by Equation (8)46: \u03b8= 360 n\u00d7NC \u00d7 S \u00f08\u00de In the equation, S represents stator slot number, NCis the period of cogging torque and calculated by: NC = 2P HCF 2P,S\u00f0 \u00de \u00f09\u00de where HCF is the highest common factor and 2P is the number of poles. In this case, the mechanical angle of rotation was calculated as 2.142 . The new model designed is presented in Figure 9. The change in cogging torque during two periods according to the position of the rotor is shown in Figure 10. In the analysis, the rotor was rotated at a mechanical angle of 0.25 , and cogging torque occurring at each position was determined. In the fragmented magnet structure, the cogging torque value was obtained as 467.3 mNm. As mentioned in Section 3, cogging torque periodically changes. This situation can be observed in Figure 10. When compared to the initial design, there was a decrease by 19.43% in cogging torque. In addition, the torque ripples in the initial design is 4.38% and in the fragmented magnet structure is 2.93%. When this method is applied to the generator, the output performance of the generator will be affected. The rated torque will drop, as the cogging torque.46 As a result of the transient analysis performed, the output power of the generator is determined as 2964 W. Rated torque has decreased from 75"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002400_012035-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002400_012035-Figure7-1.png",
        "caption": "Figure 7c. Evaluation of maximum displacement of fuselage structure made of carbon epoxy outer layers and flax epoxy inner layers for 7 psi internal pressure",
        "texts": [
            " Series: Materials Science and Engineering 1057 (2021) 012035 IOP Publishing doi:10.1088/1757-899X/1057/1/012035 Finally, pressure vessel structures made of carbon epoxy alone and pressure vessel structure made of carbon epoxy outer layers and flax epoxy inner layers are analysed and the results are compared. Both pressure vessels were of same size and support conditions, but with only material difference as discussed above. Analysis of fuselage panel has been presented as an application of NFC panels. Sample cases of modelling and post processing results are shown in Fig 7. Length of the structure was considered as 15 m for the analysis. Geometric model is created in the Solidworks software where skin, bulkheads, longerons, stringers are drawn individually and later assembled together. The thickness considered for various parts are: Skin - 0.005 m, Longerons - 0.250 m, Stringers - 0.125 m, Bulkhead - 0.150 m. The geometric model of the fuselage is created in Solidworks and then imported to ANSYS for meshing and analysis. For smooth transition of meshing between different parts, meshing is done with proper smoothening level and apt element size"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure1-1.png",
        "caption": "Fig. 1. Two commonly used types of trifilar torsional pendulum.",
        "texts": [
            " Trifilar Torsional Pendulum (TTP) is one kind of torsional pendulum methods for measuring inertia parameters of irregular-shaped rigid bodies with simple structure and measurement principle. Because it is independent to the shape of the measured bodies, and the measurement efficiency and accuracy are relatively high, it can be used in industries [7,8]. Traditional TTP and suspended TTP are two commonly used types of TTP. In the measurement using traditional TTP, the measured body need to be placed on the pendulum, as shown in Fig. 1(a), and it must make the COG of the measured body fall on the pendulum axis. The pendulum axis is defined as a line passing through the center of the pendulum and perpendicular to the plane of the pendulum. The suspended TTP shown in Fig. 1(b) has a universal joint which ensures the COG of the measured body falls on the pendulum axis automatically [8]. The TTP method is only suitable for measuring the inertia parameters of rigid bodies since these parameters will not change during measurement. After a complete measurement using suspended TTP, the mass, the COG location and the inertia tensor (including MOI and POI) of the measured rigid body, a total of ten parameters, can be obtained. * Corresponding author. E-mail address: mezhyin@scut"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001034_s40815-020-00869-y-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001034_s40815-020-00869-y-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of the electric vehicle model",
        "texts": [
            " Section 4 provides some simulation results. Section 5 gives the conclusion. Notations The matrix transposition and matrix inverse are represented by superscript \u2018\u2018T\u2019\u2019 and \u2018\u2018 1\u2019\u2019, respectively. The symbol \u2018\u2018diagf g\u2019\u2019is the block diagonal matrix and the notation \u2018\u2018 \u2019\u2019 stands for a symmetric term in a sym- metric block matrix. Meanwhile, the zero matrix and identity matrix are denoted by \u2018\u20180\u2019\u2019 and \u2018\u2018I\u2019\u2019, respectively. \u00bdA s is defined as A\u00fe AT . If a matrix is not specified in dimension, we can suppose that it has a suitable dimension. Figure 1 depicts the electric vehicle model. The sideslip angle and yaw rate can be represented by b and Xz, respectively. The vehicle mass is denoted by m. lf and lr are the distances from front and rear (FR) wheel axles to the center of gravity (CG), respectively. Iz stands for moment of inertia. vx and vy represent the vehicle longitudinal velocity and lateral velocity, respectively, where vx is a constant. The yaw moment Mz is written as [45] Mz \u00bc \u00f0Fxflcosdf \u00fe Fxrl\u00dels \u00f0Fxfrcosdf \u00fe Fxrr\u00dels where df and Fxi denote the steering input angle and the longitudinal force, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure10-1.png",
        "caption": "Fig. 10. Installation state diagram of the bottom surface",
        "texts": [
            " that the front side of the bird prevention device on the right side of the middle phase is affected by the V-shaped angle steel. The entire front side is divided into three parts, of which the two triangular mesh surfaces are symmetrical. In the middle is a symmetrical hexagonal mesh surface. 978-1-7281-6246-1/20/$31.00 \u00a92020 IEEE 810 Authorized licensed use limited to: University of Canberra. Downloaded on May 22,2021 at 12:08:06 UTC from IEEE Xplore. Restrictions apply. The bottom surface of the device is affected by the angle steel, and there is a gap in the middle section of one side, as shown in Fig.10, so that the bottom surface can reach the angle steel where the insulator hanging point is located. The bottom surface of the left and right sides of the device considers the height of the bottom angle steel on the cross arm, which also allows the design and installation of the bottom surface device to ignore the height of the screws that fix the angle steel structure of the tower. Because the angle steel extends to both sides of the cross arm on the side of the device, there is a gap in the area near the upper side of the side surfaces, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure7-1.png",
        "caption": "Fig. 7. Geometry of rings under pure axial displacement.",
        "texts": [
            " (9) From the geometric relationship between the steel ball and the raceway, the increment (> 0) of curvature center distance 0A is equal to the normal elastic contact deformation between the steel ball and raceway n\u03b4 . See Eq. (10). ( )' 0 0max ,0n A A\u03b4 = \u2212 . (10) When the bearing is subjected to a pure axial load, assuming that the inner ring is stationary and the outer ring moves by a distance of x\u03b4 under the load, the displacement of the ring at each ball position xi\u03b4 is equal to x\u03b4 , as shown in Fig. 7. The parameters 0A , 1d and 2d in Fig. 7 can be obtained according to Eqs. (6) and (7). The deformed raceway curvature center distance ' 0A is determined by Eq. (11), ' 2 ' 2 0 1 2A d d= + ( )22 1 2 xid d \u03b4= + + . (11) According to Eqs. (9) and (10), the displacement xi\u03b4 of the i-th steel ball is ' 0 0ni A A\u03b4 = \u2212 ( )22 1 2 0 xid d A\u03b4= + + \u2212 . (12) The equation of normal contact deformation ni\u03b4 of the i-th steel ball under which the pure ring axial displacement x\u03b4 occurs is established. When bearing\u2019s outer ring has a pure radial displacement in y direction (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002307_012011-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002307_012011-Figure1-1.png",
        "caption": "Figure 1. Model of a spur gearbox with two degrees of freedom.",
        "texts": [
            " On the other hand, it is much less complex than the EHL models, which allows one to use it at the design stage without the full data on the gearbox for the majority of the working conditions and doing additional experimental testing to select the right parameters. The accuracy of the proposed model was increased by considering the change to the length of the lever arm of friction as a result of vibration. For comparison purposes, the calculations were made also for three models of Coulomb friction with varied levels of detail and not considering the gear tooth friction. The analysis was conducted for a single-stage spur gear with two degrees of freedom \u03c6 1 and \u03c6 2 (Figure 1). The gears have a constant torque loads T1 and T2. Stiffness of meshing k(t) is time-varying and constitutes the main reason for gear vibrations. Two tooth stiffness models were analysed. The first one according to the ISO 6336 -1 standard [21], whereas the second one progresses similarly to the real process according to Cai [7]. Damping c in the meshing was assumed to be constant. Meshing friction was analysed for four variants of the friction model. Solid models of the gears subjected to analysis were made (Figure 2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002935_iemdc47953.2021.9449597-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002935_iemdc47953.2021.9449597-Figure4-1.png",
        "caption": "Fig. 4. Bearingless motor unit, at nominal motor current and 20% of suspension current in the x direction.",
        "texts": [
            " The bearingless motor is designed to minimize the force error angle, force ripple, and torque ripple while minimizing Authors would like to thank Business Finland for funding the project: \u201dEMBER High temperature high-efficiency oil-free heat pump, Decision number 1745/31/2020. Authorized licensed use limited to: California State University Fresno. Downloaded on June 30,2021 at 20:56:32 UTC from IEEE Xplore. Restrictions apply. losses between 17000 r/min and 30000 r/min. The AMBs and bearingless motors use standard inverter frames with direct switching control through FPGA [16]. The double 3-phase motor is controlled with current injection to common levitationmotoring windings [17]. Fig. 4 shows flux contour plots of the bearingless motor unit in nominal conditions. Cooling of the bearingless motors and SPM motor with AMBs is arranged by axially flowing vapor, R1233zd(E), entering from the rotor ends and exiting from the rotor in the axial AMB location. The temperature of the coolant entering the compressor is 60\u00b0C. Symmetricity of the twin bearingless rotor means that the cooling system can be made symmetric (e.g., coolant entering from the rotor ends and exiting from the middle)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001585_s11249-020-01339-0-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001585_s11249-020-01339-0-Figure4-1.png",
        "caption": "Fig. 4 Schematic representation of 3D view, contact angle measurement and dimension of the mold",
        "texts": [
            " The image of the droplet is captured using the camera (IDS UI-5480CP-M-GL GigE camera) attached to it. The monochromatic blue light provides a clear black image of the droplet. From the captured image, the contact angle \u03b8 is calculated by the built-in software. To achieve repetitive results, it is necessary to have an unvarying thickness of the grease and uniform surface during every test. For this purpose, a special mold made of polymer, consisting of several rhombus-shaped slots, is employed (see Fig.\u00a04), and a systematic procedure is followed in all measurements. The schematic representation of the mold in 3D view with grease in the slots and water droplet on the grease surface is shown in Fig.\u00a04a. The dimensions of the rhombus slots are as follows: length, 15\u00a0mm, breadth, 10\u00a0mm, and thickness, 0.5\u00a0mm (see Fig.\u00a04b). From the dimensions of the polymer mold, it can be inferred that this mold requires a small quantity of grease for testing. This allows the practitioner to perform several tests on a degraded grease sample extracted, for example, from a bearing in the field. Further, the multiple slots in the mold reduce the sample preparation as well as testing time, providing a constant grease height for all tests. The constant height of the samples allows easy identification of the base of the water droplet and for easy detection of the droplet edges for accurate measurement of contact angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001363_s12206-020-0734-y-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001363_s12206-020-0734-y-Figure3-1.png",
        "caption": "Fig. 3. (a) Schematic of the motor cross-section in the radial direction; (b) the thermal contact resistance network.",
        "texts": [
            " Copper loss results from the current flowing in the winding, and the winding resistance increases linearly with the temperature. Iron loss is a power loss of the core caused by a magnetizing force and divided into hysteresis and eddy current losses. In addition, different from the load losses, the iron loss is strongly dependent on material characteristics and has a fixed value regardless of operating conditions. In moving mechanical parts (i.e., shaft and bearings), additional heat is generated by friction, which is relatively negligible when compared with other energy loss terms. Fig. 3 schematically shows the radial motor components and the corresponding thermal resistance network. As described previously, the studied protrusion motor had ventilation holes in the rotor section and therefore two different radial heat transfer paths, as shown in Fig. 3(a), were predefined for presenting and comparing local thermofluidic characteristics (e.g., temperatures and radial heat transfer rates) across the solid and ventilation parts of the stator, respectively. In electric motors, temperature usually peaks at the stator coil, which yields the highest heat generation rate, owing to its Ohmic heating. Therefore, heat moves from the stator coil to the shaft and lateral frame in the radial direction. Furthermore, in Fig. 3(b), interfacial thermal resistances between solid motor components were included for comparing the predicted computational and measured experimental local temperatures. First, initial computational modeling of electric motors was conducted without considering interfacial thermal resistance. Next, the interfacial thermal resistance was adjusted such that local temperature differences between the model and experimental systems were confined to acceptable errors; this process will be discussed in more detail later",
            "0 kg \u22c5 s-1 range, as the protrusion height varies from 0 mm to 10 mm, whereas the maximal rotor temperature is inversely proportional to the protrusion height. The increased mass flow rate into the ventilation holes increases the convective heat transfer coefficient and reduces the thermal resistance in the radial direction toward the shaft, with relatively negligible temperature variation in the stator coil. Fig. 7(b) depicts the numerically estimated local temperature profiles along the two different heat transfer paths (i.e., PVent and PSolid in Fig. 3(a)) of the rotor in the radial direction. The local temperature along the solid heat transfer path (PSolid) exhibits a typical monotonous variation profile from the relatively high-temperature rotor magnet to the low-temperature shaft, whereas there is a significant temperature drop across the ventilation holes, owing to the circulating air effect (i.e., convective heat transfers between the solid and the protrusioninduced air). Fig. 8(a) shows the induced mass flow rate into the ventilation holes and the corresponding thermal characteristics, as a function of the protrusion curvature, for the protrusion height of 10 mm and zero inclination, at 3000 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001863_s12239-020-0147-z-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001863_s12239-020-0147-z-Figure5-1.png",
        "caption": "Figure 5. Piston set-liner model.",
        "texts": [
            " Where h stands for the thickness of lubrication film, and is divided into two portions: one part considered as constant corresponding to the oil film thickness between the piston and liner, and another corresponding to the oil film thickness solved by the lubrication and friction program between the piston ring and liner. Other heat transfer boundaries (such as gas in combustion chamber and cooling medium) are set to zero heat flux, that is, the so-called adiabatic boundary conditions. 5. NUMERICAL SIMULATION AND DISCUSSION 5.1. Research Object The developed model above is applied to 6110 diesel engine. The related parameters are listed in Table 1. Figure 5 shows the geometric model of piston set-liner. The high accuracy hexahedral grid is applied to generate the computational mesh, as shown in Figure 6. The whole model consists of 109475 elements and 125904 nodes. 5.2. Numerical Solution In the above all equations, the crucial equations are the heat conduction equation (29), the average Reynolds equation (11), the energy equation (25) as well as the equilibrium equation (20). In which, the heat conduction equation (29) is solved by the finite element technique (the concrete program is generated by FEPG which is an open source software) and the Newton iteration method is applied to calculate the equilibrium equation (20)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002570_iros.2011.6048138-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002570_iros.2011.6048138-Figure1-1.png",
        "caption": "Fig. 1. Structure of the ultrasonic motor with a foil type stator",
        "texts": [
            " This motor was composed of only a rotor, coiled stator, waveguide, and vibration source. It differed from the typical ultrasonic motor in that driving vibration was input from an external vibration source through a waveguide to rotate the rotor. In this research, we develop a miniature motor with a foil type stator to realize smaller size than that of the coil stator motor. In this paper, we demonstrate the driving principle of the foil stator motor and the driving test of the developed motors. Fig.1 shows the structure of the ultrasonic motor with a foil type stator. When ultrasonic vibration is transmitted to the waveguide from the ultrasonic vibrator, the vibration changes to a traveling wave and propagate to the foil type stator through the waveguide. At that moment, friction force is generated between the stator and the rotor by the vibration of the foil type stator. The rotor is rotated by this friction force. The foil-type ultrasonic motor is found to rotate in the same direction as the direction of the traveling wave; this phenomenon does not match the general ultrasonic motor principle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002826_s12206-021-0509-0-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002826_s12206-021-0509-0-Figure1-1.png",
        "caption": "Fig. 1. CAD model of 2UPR-PRU PKM.",
        "texts": [
            " Distributions of the separated DME indices for the 2UPR-PRU PKM in different operational heights are obtained and discussed in detail. The structure of this paper is as follows: Sec. 2 simply describes the mechanical structure, mobility, and inverse kinematics of the 2UPR-PRU PKM. Sec. 3 analyzes the kinematic singularity. Sec. 4 develops the inverse dynamic model of the 2UPR-PRU PKM using the screw theory and principle of virtual work. Sec. 5 discusses the dynamic performance in detail. Sec. 6 summarizes the conclusions. 2. Description, mobility, and inverse kinematics of 2UPR-PRU PKM 2.1 Structure description Fig. 1 shows the computer-aided design (CAD) model, which is composed of a fixed base, three limited-DOF limbs (two identical UPR limbs 1 1B A and 2 2B A , and PRU limb 3 3B A ), and a moving platform. These limbs have geometric relationships between axes of U joints: in the two UPR limbs, the first rotational axes of U joints coincide with each other, and the second ones are parallel with that of R joints in these two limbs; in the PRU limb, the first rotational axis of U joint is parallel to that of R joint, and the second one is parallel to that of the R joints in the other two limbs",
            " (6) Due to geometrical constraints, the position vector of origin o relative to O-xyz can be expressed as T tan 0\u23a1 \u23a4= \u23a3 \u23a6o oz z\u03b2p , thus the dependent coordinates of this PKM can be defined as: 0 tan 0 \u23a7 = \u23aa =\u23a8 \u23aa =\u23a9 ox z y \u03b2 \u03b1 (7) where \u03b1 represents the rotation angle around the z-axis. Additionally, the position vectors of ioA and iOB relative to the O-xyz can be expressed as: T T 1 1 1 2 T T 2 1 2 2 T T 3 3 3 3 0 0 , 0 0 0 0 , 0 0 0 0 , 0 0 \u23a7 = =\u23a1 \u23a4 \u23a1 \u23a4\u23a3 \u23a6 \u23a3 \u23a6\u23aa \u23aa = \u2212 = \u2212\u23a1 \u23a4 \u23a1 \u23a4\u23a8 \u23a3 \u23a6 \u23a3 \u23a6 \u23aa = \u2212 = \u2212\u23a1 \u23a4 \u23a1 \u23a4\u23aa \u23a3 \u23a6 \u23a3 \u23a6\u23a9 O o O o O o l l l l l q a R b a R b a R b . (8) Substituting Eqs. (6) and (8) into the closed-loop relationship = + \u2212i i ic p a b , as shown in Fig. 1, where ic denotes the position vector of i iB A , the expressions of the input variables of 2UPR-PRU PKM can be obtained as: ( ) ( ) ( ) ( ) ( ) 2 2 1 1 1 2 2 2 2 1 2 1 22 3 3 3 sec s c sec s c s c tan \u23a7 = + + \u2212\u23aa \u23aa\u23aa = \u2212 + \u2212\u23a8 \u23aa \u23aa = \u2212 + + \u2212\u23aa\u23a9 o o o o q z l l l q z l l l q l z l l z \u03b2 \u03b3 \u03b3 \u03b2 \u03b3 \u03b3 \u03b2 \u03b2 \u03b2 . (9) 3. Velocity and singularity analysis 3.1 Velocity analysis The relationship between the velocities of actuated joints T 1 2 3= \u23a1 \u23a4\u23a3 \u23a6q q qq and independent generalized coordinates T \u23a1 \u23a4= \u23a3 \u23a6oz\u03b2 \u03b3\u03b7 can be obtained by differentiating both sides of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0003200_s00419-021-01997-z-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003200_s00419-021-01997-z-Figure1-1.png",
        "caption": "Fig. 1 Deformation of cross-section in shear deformation element",
        "texts": [
            " However, the position vector difference function described by this method is linearly represented by the gradient vector and the coordinate, so it cannot represent the warpage deformation of the beam element [50]. For large deployable structures, the torsion of bars around the axis is inevitable due to the large size and slenderness ratio, and it is necessary to consider the warping effect in the beam element. The change of cross-section before and after deformation of beam element is shown in Fig. 1, where the beam element is composed of three nodes i , j and k. In addition, ri , r j and rk represent the positions of these three nodes in the global coordinate system, and r j,y and n j represent the gradient vector and normal direction of the section at the node j , respectively. The displacement of any point in the beam element is defined by r [ a0 + a1x + a2y + a3x ( y + 3y3 ) + a4x 2 + a5x 2y b0 + b1x + b2y + b3x ( y + 3y3 ) + b4x 2 + b5x 2y ] S(x, y)e(t) (1) where ai and bi (i 0, 1 . . "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002095_012032-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002095_012032-Figure9-1.png",
        "caption": "Figure 9. Distribution of the vertical displacement U2 in the direction of the axis OY in the case of (a) composite laminate-1, (b) composite laminate-2.",
        "texts": [],
        "surrounding_texts": [
            "It is shown that the distribution of stresses in tension test in the direction of the axis OY of both laminate specimens as shown in figure 11. 3rd International Congress on Advances in Mechanical Sciences IOP Conf. Series: Materials Science and Engineering 998 (2020) 012032 IOP Publishing doi:10.1088/1757-899X/998/1/012032 (b) laminate-2."
        ]
    },
    {
        "image_filename": "designv11_63_0000264_icems.2019.8921522-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000264_icems.2019.8921522-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram showing the distribution of the magnetic potential of stator and rotor.",
        "texts": [
            " It is enough to investigate only the forces on the three teeth due to the symmetry results from the number of slots per poles per phase (q =3) in the studied case. Alternatively, for the sake of reducing the simulation time and getting quick results, the analytical models could be used to compute the electromagnetic excitations as in [13] to [15]. In this paper, a brief overview of the analytical computations of the electromagnetic forces has been highlighted. In the following subsections, the air-gap flux density and hence the radial forces computations have been presented. Fig. 3 presents a sketch of the distribution of the scalar magnetic potential of stator and rotor respectively, from where it is readily calculating the air-gap flux density by introducing the equivalent magnetic circuit as shown in Fig. 3 [16]. In Fig. 4, and are the magnetic potential of stator and rotor respectively; is the air-gap flux density; while and represents the position of rotor position and the coordinate angle in the rotating d-q reference frame, respectively, and they agree with (3) from where is the coordinate angle in the stator stationary reference frame. In Fig. 4, \u2205 is the flux flowing through the air-gap, while \u2205 is the flux flowing through the flux barrier; and is the magnetic reluctance of the air-gap, while is the magnetic reluctance of flux barrier"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001673_icuas48674.2020.9214061-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001673_icuas48674.2020.9214061-Figure3-1.png",
        "caption": "Fig. 3. Tilt mechanism of wing-motor.",
        "texts": [
            " Motion variables notation. The angular relationships used to describe attitude and mobile aerodynamics may be generalized to describe the angular orientation of one set of axes with respect to another. For transforming the body coordinate system to the inertial system, the rotation matrix RB\u2192E is expressed, where s\u2217 = sin(\u2217) and c\u2217= cos(\u2217). RB\u2192E =  c\u03b8c\u03c8 c\u03b8s\u03c8 \u2212s\u03b8 s\u03c6s\u03b8c\u03c8\u2212 c\u03c6s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 s\u03c6c\u03b8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03c6s\u03b8s\u03c8\u2212 s\u03c6c\u03c8 c\u03c6c\u03b8  (1) For transforming the mobile aerodynamic reference to a body coordinate system (Figure 3), the rotation matrix RW\u2192B is defined, where \u00b5 is the tilt angle of the tilt mechanism of wing and rotor. RW\u2192B = c\u00b5 0 \u2212s\u00b5 0 1 0 s\u00b5 0 c\u00b5  (2) The nonlinear motion equations for the complete dynamic model of the UAV are expressed as translational motion based on Newton\u2019s second law respect to inertial system and rotational motion based on Euler\u2019s law respect to body system. m\u0304v\u0307E = RB\u2192EFB (3) I\u2126\u0307+\u2126\u00d7 I\u2126 = \u0393 B (4) Where the terms FB \u2208R3 and \u0393B \u2208R3 are the forces and moments acting on the UAV, m\u0304 = diag(m) \u2208 R3\u00d73 denotes the mass, I \u2208R3\u00d73 denotes the vehicle\u2019s inertia matrix, v\u0307E = q\u03081 = [x\u0308 y\u0308 z\u0308]T = [u\u0307 v\u0307 w\u0307]T \u2208R3 represent the linear acceleration along the body axes, \u2126 = Rq\u03072 = [p q r]T \u2208 R3 denotes the angular velocity and \u2126\u0307 = Rq\u03082 \u2208 R3 the angular acceleration according to the Euler angles around each body axis, with R as the angular velocity matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001885_icem49940.2020.9270870-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001885_icem49940.2020.9270870-Figure1-1.png",
        "caption": "Fig. 1. Perspective view of test linear motor for one armature.",
        "texts": [
            " As for control to track the resonant frequency, various research institutions and universities have been studying approaches in terms of resonant frequency control and stroke control by using, for example, waveform fitting and nonlinear observers [5] \u2013 [11]. In this paper, a position sensor-less method for estimating the resonant frequency is proposed. The frequency is estimated on the basis of the detected motor phase current, and the position is estimated using only the fundamental component of the current after removing the offset of the detected current with a high-pass filter. Simulation and experimental results show that the proposed method can successfully track the resonant frequency dynamically without a position sensor. Fig. 1 shows a perspective view of one armature of a prototype linear motor, and Fig. 2 shows a cross-sectional view taken on a plane along the A-B-C-D line in Fig. 1. The test linear motor is a single-phase machine, and the armature has a structure in which a mover having permanent magnets is sandwiched vertically. This structure, in which the armatures are arranged facing each other, has a feature in that the magnetic attractive force is canceled [12]. Moreover, since each magnetic pole has an independent structure, the magnetic pole length and the pole pitch can be easily changed. Fig. 3 shows the configuration of a test linear compressor having two armatures, and Table I shows the specifications",
            " A piston connected to the mover draws refrigerant or air into a cylinder and performs compression and discharge. The prototype linear compressor has two armatures, shown in Position Sensor-less Resonant Frequency Estimation Method for Linear Compressor with Assist Springs Takahiro Suzuki, Masaki Koyama, Shuhei Nagata, Wataru Hatsuse, Masatsugu Takemoto, and Satoshi Ogasawara A Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 23,2020 at 02:18:49 UTC from IEEE Xplore. Restrictions apply. Fig. 1, at the center, and it is driven by a single-phase inverter. Also, a linear encoder is mounted for validating the mover position opposite the cylinder side. When operating as a compressor, the linear is mounted in a sealed chamber, but this paper focuses on drive control; thus, experiments were conducted with only the motor. The motion equation of the linear motor with assist springs is defined as , (1) where M is the mover mass (kg), c is the damping coefficient (N/(m/s)), k is the spring constant (N/mm), \u03b1 is the motor thrust constant (N/A), Pc is the cylinder inside pressure (kPa), Ps is the cylinder outside pressure (kPa), S is the cross section of the piston (mm2), Fc is the detent force (N), Fs is the friction force (N), x is the mover position (mm), and Im is the motor current (A)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000631_iscmi47871.2019.9004295-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000631_iscmi47871.2019.9004295-Figure2-1.png",
        "caption": "Figure 2. Silicone rubber actuators designed for rotary motion (a) Single Rotary Actuator (b) Rotary Actuator as a pair. Parameters for these actuators are specified in Table I. These actuators have their entire body casted with soft silicone while the layers separating each air channel were made inextensible by embedding with paper inside.",
        "texts": [
            " FABRICATION OF ROTARY ACTUATORS Silicone rotary actuators were fabricated by a 2-part moulding process with the moulds built by free form fabrication technique (3D printing). The general sequence for moulding with silicone rubber are: (i) Pattern or mould making; (ii) Mould preparation; (iii) Mould filling with liquid silicone rubber; (iv) Curing; and (v) Removal of cured part. Soft rotary actuators change in angle when inflated. Because their fabrication demands casting the actuator body as a single piece, their entire body were casted with soft silicone while the layers separating each air channel was made inextensible by embedding with paper inside (Figure 2). The geometry variables for the rotary actuators (Figure 1) are given by: big radius, R1, Small radius, R2, offset distance from center of circle, d, height of air channel, h, thickness of wall, t. The geometric equations describing the area, surface area and volume are described by the following equations: ( ) ( ) ( ) ( ) ( ) ( ) ( ) The actuator consists of several compartments arranged around the circumference of a circle that can be actuated using both positive and negative pressure. Table I shows the geometric values of the rotary actuators produced in the work"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001968_me49197.2020.9286694-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001968_me49197.2020.9286694-Figure1-1.png",
        "caption": "Fig. 1. UAV considered in this study with body (B) and inertial (I) reference frames",
        "texts": [
            " Finally, a decoupled dynamic model is proposed for a 3-DOFs manipulator, and the simulation results for a real pick and place operation are presented and discussed. The paper is organized as follows: Section II describes the simulated system, Section III introduces the UAV dynamics and control, Section IV and V describe the dynamic models and report the simulation results, and Section VI concludes the paper. II. SIMULATED SYSTEM The system under consideration is composed of an octocopter UAV (Fig. 1 and Table I) and a 3-DOFs robot manipulator (Fig. 2 and Table II). A 1-DOF manipulator (Table III) is used for comparison purposes in Section IV. The UAV is a DJI S1000 heavy payload octocopter. It is composed of a carbon fiber reinforced plastics (CFRP) frame with eight arms carrying the motors and propellers, a flange for the battery that is also used as mechanical interface to the robot manipulator, and a landing gear including two retractable legs. In this study, the frame of the UAV, the propellers, and the manipulator\u2019s links are considered as rigid bodies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.3-1.png",
        "caption": "Fig. 57.3 Geometry Import and Clean-up, a Imprecise Geometry highlighted using the Geometric Diagnostics tools; b Rectification of imprecise geometry using the Geometry Edit tools",
        "texts": [
            " It was observed that Carbon + Flax + Epoxy exhibited superior properties compared to other laminates. The tensile properties were estimated both along and across the lay direction of natural fibers. Details of the mechanical properties of the laminates are available in Refs. [2] and [3]. The geometric model of the seat base was exported to ABAQUS\u00ae for finite element analysis. As the existing seat base has a contoured profile, it is prone to surface defects while exporting the files from CAD software (Fig. 57.3a). Geometric diagnostics tools were used to identify any imprecise geometry. The topology of the component was checked for smooth meshing of the component using four different criteria: free edges, solid shells, shell faces, and wire edges. As a standard guideline, smaller geometries (edges shorter than 0.1 mm and faces with areas smaller than 1 mm2) were merged with adjacent faces. Geometry edit tools were used to rectify/eliminate the imprecise geometry in the component. Figure 57.3b presents the model after clean-up using edit tools. The isotropic model approach was chosen to assign the resultant laminate properties such as Young\u2019s modulus and Poisson\u2019s ratio obtained from the experiments. A global seed size of 5 mm (Fig. 57.4a) was used for the rider seating region of the component, and seed size of 3 mm was used for rubber cushion mounting holes and the guide for rubber cushion (Fig. 57.4b). This seed size was varied depending on the thickness of the component. In order to arrive at the boundary conditions, a seat load impression test was carried out using blue print method on existing seat base under the load conditions of: nopassenger, solo passenger, and twin passenger configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001121_s00542-020-04910-w-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001121_s00542-020-04910-w-Figure2-1.png",
        "caption": "Fig. 2 Concept of the proposed MRF clutch\u2013brake activated by permanent magnets",
        "texts": [
            " Therefore, the present work aims to define through simulation, the evolution of the transmitted torque with the angular velocity for a basic single plate MRF clutch\u2013 brake. The concept has been very briefly described by Binyet and Chang (2018); its main advandtages are: simplicity in design, operation and maintenance as well as cost-effectiveness. Indeed, the permanent magnets are surrounding the clutch and triggering the MRF from the outside. The magnets can be placed in a casing that can axially slide; this allows a good shielding from the magnetic flux in off mode. The concept is shown in the schematic illustration in Fig. 2 with dimensions listed in Table 1. The magnetorheological fluid (MRF) is composed of a carrier oil and ferroparticles which can be magnetized to alter the MRF properties as schematized in Fig. 3. When there is no magnetic field applied, the ferroparticles are randomly distributed within the fluid. The behavior of the fluid can be described by the viscosity of the carrier oil. When the MRF is under the influence of a magnetic flux, the ferroparticles align in the direction of the magnetic field flux lines"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002570_iros.2011.6048138-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002570_iros.2011.6048138-Figure12-1.png",
        "caption": "Fig. 12. Experimental setup of rotational speed measurement",
        "texts": [
            " But the foil type ultrasonic motor does not have a specific preload mechanism; the clearance between rotor and stator is an important factor for adjusting the preload. To investigate the optimal clearance for high-speed rotation, the motor drive experiment is performed. The outer diameter of the stator is set to a constant 0.88 mm, while the inner diameter of the rotor varies from 0.88 mm to 0.93 mm. The rotors are prepared using an electrical discharge machine and the dimensions of the formed rotors are checked by using a laser microscope. The length of the rotors is 5 mm. Fig. 12 shows the experimental setup of rotational speed measurement. Commercially available rotary encoders are too big to measure this motor. To measure the motor rotation without creating any friction loss, non-contact measurement method is necessary. Light-weight aluminum gear is attached on the rotor, and rotation of the motor is measured by a laser displacement sensor (LKG-30, Keyence). Table V shows the result of the rotation speed measurement. It is clarified that the clearance between the rotor and the stator affects the rotation speed of the motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000032_chicc.2019.8866066-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000032_chicc.2019.8866066-Figure2-1.png",
        "caption": "Fig. 2: Schematic diagram of wheel dynamic model",
        "texts": [
            "Use the following differential equations to describe vehicle motion in the lane: Vx = (Fxfl+Fxfr) sin \u03b4+(Fyfl+Fyfr) cos \u03b4+Fyrl+Fyrr m(\u03b2\u0307+\u03b3) (1) Vy = mV\u0307x\u2212[(Fxfl+Fxfr) cos \u03b4\u2212(Fyfl+Fyfr) sin \u03b4+Fxrl+Fxrr] \u03b3 (2) \u03b3\u0307 = [(Fxfl+Fxfr) sin \u03b4+(Fyfl+Fyfr) cos \u03b4]lf Iz \u2212 ( Fyrl + Fyrr ) lr +Mz (3) Mz = tr 2 (Fxrr \u2212 Fxrl) + tr 2 ( Fxfr \u2212 Fxfl ) cos \u03b4 (4) The wheel sideslip angles are calculated as: \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b1fl = \u03b4 \u2212 ( Vy+lfr Vx\u2212 tw1 2 r ) \u03b1fr = \u03b4 \u2212 ( Vy+lfr Vx+ tw1 2 r ) \u03b1rl = \u2212 ( Vy\u2212lrr Vx\u2212 tw2 2 r ) \u03b1rr = \u2212 ( Vy\u2212lrr Vx+ tw2 2 r ) (5) The wheel dynamics is shown in Fig.2 and characteristiced by Eq.(6),where \u03c9i(i=fl,rl,fr,rr) are wheel rotational velocities,R is tire effective radius,J\u03c9 is wheel moment of intertia,and Ti is motor torque(Ti > 0 for motor driving torque and Ti < 0 for motor regenerative braking torque). \u03c9\u0307 = 1 J\u03c9 (\u2212RwFxi + Ti) (6) Tire lateral force is related to various factors such as road surface adhesion,load,speed and tire material.Analytical tire models (GIM models) and semi-empirical models (magic formulas),which are widely used in the literature,can accurately calculate the lateral force of tire on the specific attached road surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001517_tmech.2020.3021905-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001517_tmech.2020.3021905-Figure2-1.png",
        "caption": "Figure 2. The goal is to drive the robot end-effector to the target position and control \u03c5cur (k) to be aligned with \u03c5tar.",
        "texts": [
            " However, without considering the nonlinear structure of SO (3), it is difficult to express the optimization objective as a closed-form function of the IOC, especially when the system model is unknown. Thus the timeconsuming searching procedure makes the model-less control impractical in application. Moreover, all the aforementioned methods cannot be applied to 4-DOF or 5-DOF manipulators, which are not functionally redundant. Let us define the following notations at discrete time k: \u2022 pcur (k): The current position of the end-effector. \u2022 Rcur (k): The current orientation of the end-effector. As shown in Fig. 2, the rotation around the given direction has no effect on the task performance. Therefore, only three position DOFs and two orientation DOFs are required for a given target, namely dim(T ) = 5. Without Authorized licensed use limited to: Cornell University Library. Downloaded on September 09,2020 at 21:12:30 UTC from IEEE Xplore. Restrictions apply. 1083-4435 (c) 2020 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_63_0003246_ldia49489.2021.9505885-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003246_ldia49489.2021.9505885-Figure1-1.png",
        "caption": "Fig. 1. The structure of a linear synchronous motor",
        "texts": [
            " It can be seen that the present research on linear synchronous motor mostly considers the ideal modeling for the purposes of investigating the electromagnetic field, control method, faults and so on. In order to consider the actual installation and power supply process of linear synchronous motor, this paper establishes the field-circuit coupling model under the condition of inverter power supply based on 600km/h operation condition, and the internal electromagnetic field, traction force and levitation force of the motor are calculated. Then, the influence of parameter and inverter faults on the traction and levitation force of linear synchronous motor is analyzed. VERIFICATION Fig. 1 shows the schematic diagram of the electromagnet module supported on one side of the single carriage in the middle of the maglev train. The long stator is designed in sections with each interval of 5mm, and the calculation model includes nine stator segments. The mover is distributed symmetrically including ten complete poles and 20 21 1 3t h In te rn at io na l S ym po siu m o n Li ne ar D riv es fo r I nd us tr y Ap pl ic at io ns (L DI A) | 9 78 -1 -7 28 1- 72 10 -1 /2 1/ $3 1. 00 \u00a9 20 21 IE EE | D O I: 10 ",
            " Authorized licensed use limited to: University of Gothenburg. Downloaded on September 01,2021 at 02:40:32 UTC from IEEE Xplore. Restrictions apply. two end poles, and the top of complete poles are slotted and evenly embedded in the generator windings to form a linear generator system. The stator core, mover core and magnetic yoke are all laminated with silicon steel sheets, and the windings are all made of aluminum wire. The basic parameters corresponding to the model are shown in Table I . According to Fig. 1, the structure of the linear synchronous motor investigated in this paper is highly symmetrical, and the two-dimensional finite element method can meet the calculation requirements. In order to simplify the solution process, the following assumptions are considered. (1) It is assumed that the permeability of the stator core is isotropic. (2) It is assumed that the current in the conductor is uniformly distributed. (3) There is no free charge in the studied field. Based on the assumptions and the electromagnetic calculation equation, the corresponding two-dimensional transient electromagnetic boundary value equation is as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002509_14644193211003777-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002509_14644193211003777-Figure4-1.png",
        "caption": "Figure 4. Exaggerated view of vector definition of rolling element relative to bearing raceway.",
        "texts": [
            " The differential equation of the disk is shown in equation (10) mdiski\u20acxdiski \u00fe 3EI L3 i 2xdiski xds xfs\u00f0 \u00de \u00bc 0 mdiski\u20acydiski \u00fe 3EI L3 i 2ydiski yds yfs\u00f0 \u00de \u00bc 0 Idiski;t\u20achdiski;X \u00fe Ipxdisk _hdiski;Y \u00fe 3EI L3 i 2L2hdiski;X Liyds \u00fe Liyfs \u00bc 0 Idiski;t\u20achdiski;Y \u00fe Ipxdisk _hdiski;X \u00fe 3EI L3 i 2L2hdiski;Y Lixds \u00fe Lixfs \u00bc 0 8>>>>>>< >>>>>>: (10) where \u20acxdiski and \u20acydiski (i\u00bc 1, 2, . . . , n) are the acceleration of the i-th disk in the X direction and Y direction, \u20achdiski;X and \u20achdiski;Y are the angular acceleration of the ith disk rotated in the X direction and Y direction. Differential equation of motion for bearing In order to define the energies and generalized forces, the location of the rolling element relative to the inner and outer raceways can be described by the position vectors of the rolling element, which is shown in Figure 4. In the Figure 4, rs and rh are the position vector of the inner and outer raceway\u2019s geometrical center relative to the inertial reference frame respectively, qi is the position of the i-th rolling element relative to the inner raceway, mi is the mass of the i-th rolling element, r is the radius of inner raceway, R is the radius of outer raceway, Co is the center of the outer raceway, Ci is the center of the inner raceway, hi and hxi are the angular position of the i-th rolling element relative to the inner and outer raceway\u2019s Xaxis respectively The kinetic energy of each rolling element Ti is given by equation (11)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001528_0954407020951318-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001528_0954407020951318-Figure2-1.png",
        "caption": "Figure 2. Two-stage stiffness arc-spring structure.",
        "texts": [
            " The component force along the radial direction of the distribution of the arc-spring increases with the increases of the twist angle between the first mass and the second mass, thus, the friction increases as the twist angle increases. This section will analyze the stiffness and damping of the DMF. The DMF stiffness is generated by a two-stage piecewise arc-spring structure, which is composed of an outer arc-spring with small stiffness (angular stiffness is ka1) and an inner arc-spring with large stiffness (angular stiffness is ka2). 13 The structure of the two-stage piecewise arc-spring is shown in Figure 2. The inner arcspring is nested in the outer arc-spring. When the springs are in the free state, the distribution gap between the inner and outer arc-spring is u1. When the engine is working under idle speed or light load condition, the external arc-spring compression deformation is less than u1, thus, only the outer arcspring is involved in the work of the DMF, and the first stage stiffness of DMF is: K= ka1 When the engine is working under heavy load conditions, the external arc-spring compression deformation is more than u1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001672_s00231-020-02959-x-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001672_s00231-020-02959-x-Figure1-1.png",
        "caption": "Fig. 1 Two dimensional schematic of the journal bearing",
        "texts": [
            " This code is developed based on direct heat transfer using the finite volume method. For validation of prepared codes, the results are compared with another experimental data. By this research the optimum number of sensors and their positions are obtained. Proper time step for temperature measurements and the required numbers of future time steps for inverse method are presented. In this paper, a two dimensional model of a pressure-fed journal bearing is investigated thermally. It is composed of three main parts, journal, bearing and lubricating oil (Fig. 1). The validation and subsequent results are obtained for a specified geometry (Table 1), previously used by Kucinschi et al. [10]. Several sensors are considered in the bearing to measure the temperature. These data are used by inverse code to obtain the heat flux distribution on the inner surface of the bearing. The number of sensors is considered eight as default values (Fig. 2). The sensors are situated at the same distance to neighboring sensors in the circumferential direction, but their radial positions should be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002105_j.measurement.2020.108956-Figure16-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002105_j.measurement.2020.108956-Figure16-1.png",
        "caption": "Fig. 16. Non coincidence error of X-axes of two RCSs.",
        "texts": [
            " The defined transmission RCS is as shown in Fig. 15 (b). Definition of powertrain RCS: powertrain RCS is usually defined as the same one as engine RCS as shown in Fig. 15 (c). The factors influencing assembling errors for inertia parameters of powertrain are: (1) The X-axis of engine RCS does not coincide with X-axis of the transmission RCS. The two X-axes of the RCS are theoretically coincident, but since they are measured separately and due to the measurement errors, they are not coincident in reality, as shown in Fig. 16. Since there is measured error for the measured X-axis, errors also existed for Y-axis and Z-axis, which will cause errors in rotation transformation of the inertia parameters of the transmission, and thus cause the resulting inertia parameters to deviate from the real. The definitions of symbols in Fig. 16 are shown in Table 17. T. Li et al. Measurement 174 (2021) 108956 (2) The origin point of the transmission RCS and engine RCS do not coincide. Since the two X-axes for RCSs of the engine and the transmission do not coincide, the intersection points between the two X-axes with the contacting face between the engine and the transmission are not coincide, as shown in Fig. 17. The distance between the two origin points can be measured only after the engine and transmission are assembled, but in reality the distance cannot be measured because the RCSs are inside the powertrain"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001551_ijbic.2020.109715-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001551_ijbic.2020.109715-Figure3-1.png",
        "caption": "Figure 3 CAD design of the simplified wing\u2019s vein location as well as the critical locations, (a) the forewing",
        "texts": [
            " The main vein structures were manually traced out and divided into sections using splines and polygons. The smaller vein structures were then modelled, in a similar way, to match the detailed patterns. The spatial network analysis method was employed to perform the segmentations based on the pattern density value in order to biomimic the actual wing of a dragonfly (Sivasankaran and Ward, 2016). An initial wire frame, two-dimensional scaled model was constructed using SolidWorks to create a three-dimensional solid model by extruding the imported two-dimensional wire frame as shown in Figure 3. 3D printing (rapid prototyping or additive manufacturing) is a fabrication method that allows the creation of 3D objects from digital sources. The MakerBot 2X 3D printer used in this study is depicted in Figure 4. The obtained CAD design illustrated in Figure 3 then transferred into 3D printable form with the use of a MakerBot desktop application. A 3D object is drawn by placing successive layers of materials to create the object. Each layer is a thin cross-section of the object. Layers are put down using fused deposition modelling (FDM), in which a coil of plastic filament is unwound into a heated extrusion head (nozzle) illustrated in Figure 4. The nozzle is heated to a high temperature in order to melt the filament material, as the nozzle moves (vertically and horizontally) in a pattern according to the digital model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001440_s00170-020-05863-0-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001440_s00170-020-05863-0-Figure5-1.png",
        "caption": "Fig. 5 Contour plot of temperature distribution results at the starting point",
        "texts": [],
        "surrounding_texts": [
            "The increased speed of the reinforcement particles that displays a proper flow of molten liquid is brought about by the increase in laser power; also noting that an increased laser power causes improvement, standardized distribution, and spheroidization as shown in Figures 3, 4, and 5. As the laser beam moves from one point to another as shown in Figs. 3, 4, and 5, the temperature changes and the microstructures are determined by the laser input, scanning speed, and fast cooling rate. Figure 6 shows the microstructure of the titanium alloy base metal which acts as heat sink. High cooling rates form dendritic structures within the coatings. The major factors that determine the formation of the dendritic structure are the thermal gradients within the substrate during cooling and the cooling rates as shown in Fig. 7. The faster and slower cooling rates cause the spacing of the dendritic arm. Dendrite arm spacing has great influence on the mechanical properties of additive manufactured parts. The dendrite arm spacing is divided into primary and secondary. Crystal structures are modified by the rate of solidification in the microstructure. The distance from the Ti-6AI-4V alloy substrate determines the peak temperature distribution in the molten pool. There is a notable affiliation between the powder feed rate and laser scanning speed. The turbulent Marangoni convection current is the momentum and turbulence that occurs in the melt pool, which in fact is brought about by the blowing of the powders into the melt pool through the argon gas which carries the powders during laser metal deposition [20, 21, 26]. A decrease in the aspect ratio of columnar grains comes into effect by the increase in equiaxed grains as the build height increases. Processing conditions on grain structure with its influence were also noted. The microstructural development can also be affected by other processing parameters such as the powder flow rate as suggested by the results. Therefore, the microstructural development may have been conclusively determined by the input energy density, which could be defined by laser power, scanning speed, powder flow rate, and thermal history. Hence, there is influence on microstructure when any processing parameter affects the energy density and thermal history. The increase in line energies reduces its influence on the molten pool width, as confirmed by the simulations results. Only for various process parameter combinations were the molten pool depth determined by simulations, with constant line energy resulting in near constant melt pool depth. Higher line energies increased the molten pool depth. The temperature history within the material is prevailed by the process, which influences the mechanical properties of the fabricated coatings."
        ]
    },
    {
        "image_filename": "designv11_63_0000496_oceans40490.2019.8962582-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000496_oceans40490.2019.8962582-Figure2-1.png",
        "caption": "Fig. 2. Simulated reflection (solid blue in online version) and transmission (dashed red in online version) characteristics for the reconfigurable voodoo-",
        "texts": [
            ", parallel to the electric field ), currents are induced along the pins, resulting in a strong interaction with propagating waveguide structure with pins filling the holes. (a) and . (b) and . fields. This would not be the case if the pins were placed horizontally. Simple single- and double-pin discontinuities located in the -plane have been studied analytically [12], [13]. As an example, the reflection and transmission characteristics of the reconfigurable voodoo-waveguide structure incorporating two pins is shown in Fig. 2 for two different pin combinations. Polar plot responses exhibit a spiral frequency response (for both reflection and transmission measurements), which is unique to the reconfigurable voodoo structure. This demonstrates that this single-component verification kit should be sufficient for most practical applications. Having values spread out across most of the scattering ( )-parameter planes provides a comprehensive verification for the VNA\u2019s reflection and transmission measurements, as will be discussed in detail in the following sections",
            " Reconfigurable Structure Electrical Measurements The voltage-wave reflection and transmission coefficients of the reconfigurable voodoo-waveguide structure were first measured in order to obtain the reference data for the device. With our design, which has 32 different pin combinations, not all combinations had sufficiently distinct and useful characteristics. Some of the most illustrative combinations are shown in Figs. 5 and 6. As can be seen, the reconfigurable nature of the structure allows a diverse range of reflection and transmission values. The same behavior is also observed as spirals in the measurement complex planes, as shown in Fig. 2. This diverse range has the advantage that most parts of the -parameter complex planes can be verified, and hence, more detailed information about the quality of calibration can be obtained. This contrasts with the behavior of most commercial verification kit components (i.e., attenuators and waveguide sections) that only provide relatively flat magnitude frequency responses (corresponding to circles in the measurement complex planes). In order to verify the repeatability of the reconfigurable voodoo-waveguide device, two independent studies were undertaken in order to decompose the random errors introduced by flange misalignments and pin insertion misalignments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000510_tensymp46218.2019.8971311-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000510_tensymp46218.2019.8971311-Figure2-1.png",
        "caption": "Figure 2 shows the control strategy. Speed error is given to a PI controller to generate the reference torque Te *. The reference angle between stator and rotor flux vectors *",
        "texts": [
            " From the figure we can write, cos sin cos cos ds s fs qs s fs dr r fr qr r fr \u03b8 \u03b8 \u03b8 \u03b8 \u03a8 =\u03a8 \u03a8 =\u03a8 \u03a8 =\u03a8 \u03a8 =\u03a8 (3) Substituting equation (3) in equation (1). 3 ( sin cos cos sin ) 2 2 m s r fs fr s r fs fr s r LPT L L \u03b8 \u03b8 \u03b8 \u03b8 \u03c3 = \u03a8 \u03a8 \u2212 \u03a8 \u03a8 3 sin( ) 2 2 3 ( sin ) 2 2 m s r fs fr s r m s r sr s r LP L L LP L L \u03b8 \u03b8 \u03c3 \u03b8 \u03c3 = \u03a8 \u03a8 \u2212 = \u03a8 \u03a8 (4) The equation (4) [4] is used in the proposed scheme to control the torque. sr\u03b8 is determined from the reference torque Te * using equation (5). ! \"#$ $\" # $ \"$ % ! &' ($% ) # Authorized licensed use limited to: Carleton University. Downloaded on July 26,2020 at 18:30:52 UTC from IEEE Xplore. Restrictions apply. Fig.2 Control strategy Fig.3 Flux generation logic * * Sin Te srKs r \u03b8= \u03a8 \u03a8 (5) Here, the magnitude of the reference stator flux * s\u03a8 is kept fixed at its rated value. The stator flux angle * fs\u03b8 needed to produce * sr\u03b8 is determined (See Fig.1 and Fig.3). The rotor flux angle fr\u03b8 is assumed to be a constant during that interval of time due to larger time constant in the rotor circuit. Hence, we can write, sr fs fr\u03b8 \u03b8 \u03b8= \u2212 (6) * * sr fs fr\u03b8 \u03b8 \u03b8= \u2212 (7) Using (6) and (7), we can write * * fs sr s r fs\u03b8 \u03b8 \u03b8 \u03b8= \u2212 + (8) The voltage equation of an induction motor in the stator reference frame is given by d sV i rs s s dt \u03a8 += (9) The voltage drop across stator resistance of an induction machine is very small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000924_s026357472000020x-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000924_s026357472000020x-Figure5-1.png",
        "caption": "Fig. 5. A cylindrical link with circular cross section is inclined at \u03b4\u25e6 about Y axis. The inclined link has an elliptical cross section taken parallel to XY plane. This cross-section area is used for computing the MOI.",
        "texts": [
            "org/core. San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at The MOI of inclined link is different from that of the non-inclined link because the cross-section area about the vertical axis has changed. Assuming cylindrical links (with circular cross section) for the manipulator legs, MOI for the links are determined. Let the cylindrical link be of radius Rc and length L is inclined about Y axis by an inclination angle (\u03b4) as shown in Fig. 5. In order to determine the MOI (second moment of area) about the Z axis (refer Fig. 5), the crosssection area of the inclined link in the XY plane is considered. The cross-section area is an ellipse with minor diameter equals Rc (along Y axis) and major diameter is Rc/cos\u03b4 (along X axis), as the link is inclined about an angle \u03b4. Area of cross section A = \u03c0 RC RC cos \u03b4 (1) https://www.cambridge.org/core/terms. https://doi.org/10.1017/S026357472000020X Downloaded from https://www.cambridge.org/core. San Francisco State University, on 09 Nov 2020 at 05:57:53, subject to the Cambridge Core terms of use, available at It can be noted that as \u03b4\u2192 0\u25e6, the cross-section area is a circle and as \u03b4\u2192 90\u25e6, the major diameter of ellipse RC/ cos \u03b4\u2192 \u221e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002766_j.triboint.2021.107098-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002766_j.triboint.2021.107098-Figure2-1.png",
        "caption": "Fig. 2. General overview of the procedure [4].",
        "texts": [
            " This implementation is not straightforward since some mathematical work is * Corresponding author. E-mail address: inigo.martin@ehu.eus (I. Mart\u0301In). Contents lists available at ScienceDirect Tribology International journal homepage: www.elsevier.com/locate/triboint https://doi.org/10.1016/j.triboint.2021.107098 Received 15 March 2021; Received in revised form 27 April 2021; Accepted 14 May 2021 Tribology International 161 (2021) 107098 needed but, once done, the methodology can be applied to any bearing geometry. Fig. 2 shows a general overview of previous [4] and current work, where the proposed procedure is explained. In the following sections, the analytical formulation in [4] will be briefly described as the starting point for the FE approach. Then, its implementation in the FE model together with another spring-based simplification will be explained. Finally, the results of these efficient FE models will be shown, and compared with a FE reference analysis for validation purposes. For the sake of clarity and completeness of the work, the analytical formulation in [4] will be summarized here. As mentioned before, and illustrated in the upper part of Fig. 2, the formulation reproduces the structural behaviour of the wire-roller-wire subset with an equation system. In particular, a set of 6 equations is needed, which relate the parameters shown in the representation of the analytical model in Fig. 3a. The basis of this formulation lies on assuming that the roller-wire contact remains in stick status and the wire-ring contacts in slip status. This way, the wire-roller-wire subset behaves as a rigid solid. N2 = k2\u22c5\u22062 (2) (N2 \u00b1 \u03bc\u22c5N1)\u22c5cos(\u03b1)+ (N1 \u2213 \u03bc\u22c5N2)\u22c5sin(\u03b1) = k3\u22c5\u22063 (3) DCW/2\u22c5cos(\u03b10) = \u2206R/2 \u2212 \u22062 +(DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) (4) DCW/2\u22c5sin(\u03b10) = \u2206A/2 \u2212 \u22061 +(DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) (5) N2\u22c5((DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) ) \u00b1 \u03bc\u22c5N1\u22c5(((\u03bb/2 \u2212 \u22061) + (DCW/2 \u2212 \u22063)\u22c5sin(\u03b1) ) \u00b1 \u03bc\u22c5N2\u22c5((\u03bb/2 \u2212 \u22062) + (DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) ) \u2212 N1\u22c5((DCW/2 \u2212 \u22063)\u22c5cos(\u03b1) ) = 0 (6) According to Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000901_s11012-020-01160-y-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000901_s11012-020-01160-y-Figure7-1.png",
        "caption": "Fig. 7 A 3D finite element model built in ABAQUS",
        "texts": [
            "5 Comparison with a finite element model To validate the overall deformation of crowned tooth pair derived in this paper, the deformation of the same gear pair is evaluated in this section using the finite element method for comparison. The specifications of the spur gear pair used in the proposed analytical method and FEM are presented in Table 2. In many earlier literatures [11, 16, 22], in order to simplify the calculation of tooth deflection, only one pair of teeth was considered. The same simplification is also introduced in the present study. As shown in Fig. 7, a 3D finite element model with 477,768 elements has been built in ABAQUS environment [23]. The entities used in the finite element model are established by utilizing UGNX software according to profile equations in Sect. 2.1. The gear pair in the model is meshed using the element type C3D8R. The whole model is mapped hexahedral. All the nodes of wheel hub are fixed. Every node of pinion hub is coupling with the reference point which is on the pinion axis and has only one degree of freedom rotating around the axis",
            "02% and the mean difference is 10.92%. Thus, the differences are at a relatively low level when the torque applied on the gear pair is not more than 400 N m. As the overall deformation results from the proposed analytical method and FEM show good agreements, the validity of the proposed method has been proved. Nevertheless, it is important to note that with the increasing of the applied load, the contact area of gear tooth pair becomes larger, which may lead to invalidation of the half-space assumption. From Fig. 7a\u2013d, we can also observe that the amount of total deformation increases less than twice under the doubled load, which suggests that meshing stiffness of the gear pair increases with the applied load. 3.1 Normal surface thermoelastic displacement calculation A point analysis model of thermoelastic displacements in moving bodies proposed by Liu et al. [7, 8] is introduced to calculate the coupled thermal\u2013mechanical deflection of a crowned gear pair under a meshing load. The details about the thermoelastic displacement analysis method can be found in Refs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002113_0954407020984668-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002113_0954407020984668-Figure9-1.png",
        "caption": "Figure 9. Representation of the synthesis step that reduced the stiffness elements of the theoretical model (a) to the equivalent stiffness of the simplified model (b).",
        "texts": [
            " From this observation, confirmed by other tests conducted on other examples of steering columns, the authors have defined the first hypothesis of the proposed simplified modeling procedure: the steering column inertia does not influence system dynamics and, therefore, the mass of the whole can be traced and reduced to the mass of the rack (which typically assumes values of approximately 2 kg), then assigning an unitary value to the inertia of the steering column. Considering the steering model of Figure 9(a) (repetition of that of Figure 5 but in which friction torques to the joints and the relative stiffness and rotations are represented) the equilibrium equations of the static can be written: M1 =KJ1 u1 \u00f07\u00de M2 =KJ2 u2 = M1 t12 \u00f08\u00de M3 =Kt u3 = M1 t12 t23 \u00f09\u00de with M1 applied moment to the steer, KJ1 torsional stiffness connecting the upper universal joint, u1 torsional angle associated with the steering column, M2 moment transmitted to the intermediate shaft, KJ2 torsional stiffness connecting the bottom universal joint, u2 torsional angle associated to the intermediate shaft, M3 moment transmitted to the input shaft, KT torsional stiffness representative of the torsion bar, u3 torsional angle associated with the torsion bar, and t12, t23 transmission ratios of the upper and lower universal joint, respectively",
            " In order to synthetize the system stiffness into a unique equivalent parameter KTBar, the following equations of equilibrium and congruence can be written: M1 =KTBar u \u00f010\u00de with: u= u1 + u2 t12 + u3 t12 t23 \u00f011\u00de Through appropriate substitutions: M1 KTBar = M1 KJ1 + M1 t122 KJ2 + M1 t122 t232 Kt \u00f012\u00de and so: 1 KTBar = 1 KJ1 + 1 t122 KJ2 + 1 t122 t232 Kt \u00f013\u00de It is confirmed by this analysis that the transmission ratio of universal joints is not constant. The analysis of the state of the art in terms of typical angles for steering system universal joints leads to affirm and demonstrate that the equivalent stiffness can be approximated, neglecting the ratio introduced by the universal joints themselves (14) (Figure 9(b)). As proof of this statement, referring to the case of Table 1 test (a1 =17 and a2 =11 ) it can observed how transmission ratio varies, cycling in the range 0.954t124 1.05 and 0.984t234 1.02, confirming the above statement (14). In that specific test case, the calculated value of the equivalent stiffness of the system is KTBar =1392 N mm=deg. It remains of course possible to obtain the exact value by using equation (13) avoiding the considered approximation stated by this simplification methodology expressed by (14)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002307_012011-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002307_012011-Figure6-1.png",
        "caption": "Figure 6. Determining the minimum and maximum arms of sliding friction for the pinion and a wheel. \ud835\udc5f\ud835\udc4e \u2212 tip radius, \ud835\udc5f\ud835\udc64 \u2013 working pitch radius, \ud835\udc5f\ud835\udc4f \u2013 base radius, \ud835\udc5f\ud835\udc531\ud835\udc5a\ud835\udc56\ud835\udc5b \u2013 min value of moment arm of sliding friction force acting on pinion, \ud835\udc5f\ud835\udc531\ud835\udc5a\ud835\udc4e\ud835\udc65 \u2013 max value of moment arm of sliding friction force acting on the pinion, \ud835\udc36 \u2013 pitch point, \ud835\udc34, \ud835\udc38 \u2013 points where gears start and end to be in contact.",
        "texts": [
            "1088/1742-6596/1736/1/012011 Friction in the assumed gearbox model with 2 DOF can only be considered as the moment of friction. For this reason, the arm of sliding friction is to be determined. The speed of the movement of the contact point on the line of action is constant for the constant angular speed and with uninterrupted teeth contact. It stems from the rule of evenness of the projections of the speed of points of a solid body in a mutual direction. For the pinon, the minimal arm of friction is equal to the length of the section |L1A|, whereas the maximum |L1E| (Figure 6). In order to determine the arm |L1A| and |L2E| line segment on the LOA depending on the known values will be specified: |\ud835\udc3f1\ud835\udc38| = \u221a|\ud835\udc421\ud835\udc38|2 \u2212 |\ud835\udc421\ud835\udc3f1|2 = \u221a\ud835\udc5f\ud835\udc4e1 2 \u2212 \ud835\udc5f\ud835\udc4f1 2 (16) |\ud835\udc3f1\ud835\udc36| = |\ud835\udc421\ud835\udc36| sin \ud835\udefc\ud835\udc64 = \ud835\udc5f\ud835\udc641 sin \ud835\udefc\ud835\udc64 (17) and similarly |\ud835\udc3f2\ud835\udc34| = \u221a|\ud835\udc422\ud835\udc34|2 \u2212 |\ud835\udc422\ud835\udc3f2|2 = \u221a\ud835\udc5f\ud835\udc4e2 2 \u2212 \ud835\udc5f\ud835\udc4f2 2 (18) |\ud835\udc3f2\ud835\udc36| = |\ud835\udc422\ud835\udc36| sin \ud835\udefc\ud835\udc64 = \ud835\udc5f\ud835\udc642 sin \ud835\udefc\ud835\udc64 (19) On the basis of this, the line segment for the pinion can be calculated |\ud835\udc3f1\ud835\udc34| = |\ud835\udc3f1\ud835\udc38| \u2212 (|\ud835\udc3f1\ud835\udc38| \u2212 |\ud835\udc3f1\ud835\udc36|) \u2212 (|\ud835\udc3f2\ud835\udc34| \u2212 |\ud835\udc3f2\ud835\udc36|) = |\ud835\udc3f1\ud835\udc36| \u2212 |\ud835\udc3f2\ud835\udc34| + |\ud835\udc3f2\ud835\udc36| (20) therefore \ud835\udc5f\ud835\udc531\ud835\udc5a\ud835\udc56\ud835\udc5b = |\ud835\udc3f1\ud835\udc34| = \ud835\udc5f\ud835\udc641 sin \ud835\udefc\ud835\udc64 \u2212 \u221a\ud835\udc5f\ud835\udc4e2 2 \u2212 \ud835\udc5f\ud835\udc4f2 2 + \ud835\udc5f\ud835\udc642 sin \ud835\udefc\ud835\udc64 = (\ud835\udc5f\ud835\udc641 + \ud835\udc5f\ud835\udc642) sin \ud835\udefc\ud835\udc64 \u2212 \u221a\ud835\udc5f\ud835\udc4e2 2 \u2212 \ud835\udc5f\ud835\udc4f2 2 (21) \ud835\udc5f\ud835\udc531\ud835\udc5a\ud835\udc4e\ud835\udc65 = |\ud835\udc3f1\ud835\udc38| = \u221a\ud835\udc5f\ud835\udc4e1 2 \u2212 \ud835\udc5f\ud835\udc4f1 2 (22) and for the gear |\ud835\udc3f2\ud835\udc38| = |\ud835\udc3f2\ud835\udc34| \u2212 (|\ud835\udc3f2\ud835\udc34| \u2212 |\ud835\udc3f2\ud835\udc36|) \u2212 (|\ud835\udc3f1\ud835\udc38| \u2212 |\ud835\udc3f1\ud835\udc36|) = |\ud835\udc3f2\ud835\udc36| \u2212 |\ud835\udc3f1\ud835\udc38| + |\ud835\udc3f1\ud835\udc36| (23) therefore \ud835\udc5f\ud835\udc532\ud835\udc5a\ud835\udc56\ud835\udc5b = |\ud835\udc3f2\ud835\udc38| = (\ud835\udc5f\ud835\udc641 + \ud835\udc5f\ud835\udc642) sin \ud835\udefc\ud835\udc64 \u2212 \u221a\ud835\udc5f\ud835\udc4e1 2 \u2212 \ud835\udc5f\ud835\udc4f1 2 (24) \ud835\udc5f\ud835\udc532\ud835\udc5a\ud835\udc4e\ud835\udc65 = |\ud835\udc3f2\ud835\udc34| = \u221a\ud835\udc5f\ud835\udc4e2 2 \u2212 \ud835\udc5f\ud835\udc4f2 2 (25) The change to the arm of friction is linear for involute meshing and a constant rotational speed of both wheels",
            "1088/1742-6596/1736/1/012011 and for the wheel \ud835\udc5f\ud835\udc532 = \ud835\udc61 \u2219 tan (180 \u2212 tan\u22121 ( \ud835\udc5f\ud835\udc532\ud835\udc5a\ud835\udc4e\ud835\udc65\u2212\ud835\udc5f\ud835\udc532\ud835\udc5a\ud835\udc56\ud835\udc5b \ud835\udc61\ud835\udc5d )) + \ud835\udc5f\ud835\udc532\ud835\udc5a\ud835\udc4e\ud835\udc65 (27) The results of the above-presented equations including the time sequence of the meshing of teeth are shown in Figure 8 (in Figures 8, 10-12 the blue colour means two pairs of teeth in contact and green one pair). Figure 7. Theoretical schemes of the progress of changes to the arm of the sliding friction. It is already possible to determine the Coulomb moment of friction \ud835\udc40\ud835\udc53\ud835\udc50, the equation of which for the pinion and wheel is: \ud835\udc40\ud835\udc53\ud835\udc50 = \ud835\udc39\ud835\udc53\ud835\udc5f\ud835\udc53 (28) The direction of the force and so the moment of friction depends on the relative speed. In the pitch point C, the relative sliding velocity equals 0. After crossing this point change to the sense of the friction force occurs (Figure 6). Time from entering the tooth contact to the pitch point can be determined from the following dependency: \ud835\udc61\ud835\udc36 = |\ud835\udc3f1\ud835\udc36|\u2212|\ud835\udc3f1\ud835\udc34| \ud835\udf141\ud835\udc5f\ud835\udc4f1 (\ud835\udc60) (29) CMES 2020 Journal of Physics: Conference Series 1736 (2021) 012011 IOP Publishing doi:10.1088/1742-6596/1736/1/012011 In the model of friction considering the rolling friction, information on the value of the sliding velocity is required. According to the symbols in Figure 9 linear speed in the contact point equals the product of the angular speed and radius: \ud835\udc631 = \ud835\udf141\ud835\udc45\ud835\udc631 (30) \ud835\udc632 = \ud835\udf142\ud835\udc45\ud835\udc632 (31) where: \ud835\udc45\ud835\udc631 = \u221a\ud835\udc5f\ud835\udc4f1 2 + \ud835\udc5f\ud835\udc531 2 \ud835\udc45\ud835\udc632 = \u221a\ud835\udc5f\ud835\udc4f1 2 + \ud835\udc5f\ud835\udc532 2 The value of the tangential velocity to the profile is: \ud835\udc63\ud835\udc611 = \ud835\udc631 sin \ud835\udefd1 = \ud835\udc631 sin (tan\u22121 ( \ud835\udc5f\ud835\udc531 \ud835\udc5f\ud835\udc4f1 )) (32) \ud835\udc63\ud835\udc612 = \ud835\udc632 sin \ud835\udefd2 = \ud835\udc632 sin (tan\u22121 ( \ud835\udc5f\ud835\udc532 \ud835\udc5f\ud835\udc4f2 )) (33) where: \ud835\udefd1 = tan\u22121 ( \ud835\udc5f\ud835\udc531 \ud835\udc5f\ud835\udc4f1 ) \ud835\udefd2 = tan\u22121 ( \ud835\udc5f\ud835\udc532 \ud835\udc5f\ud835\udc4f2 ) The value of sliding velocity for the pinion is equal \ud835\udc63\ud835\udc601 = \ud835\udc63\ud835\udc611 \u2212 \ud835\udc63\ud835\udc612 (34) and for the wheel \ud835\udc63\ud835\udc602 = \ud835\udc63\ud835\udc612 \u2212 \ud835\udc63\ud835\udc611 (35) CMES 2020 Journal of Physics: Conference Series 1736 (2021) 012011 IOP Publishing doi:10"
        ],
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    },
    {
        "image_filename": "designv11_63_0001399_s00170-020-05913-7-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001399_s00170-020-05913-7-Figure4-1.png",
        "caption": "Fig. 4 Forming process of asymmetric spinning using the translational pass set",
        "texts": [
            " The linear pass is used at the end of the process to push the wall on the mandrel. In asymmetric spinning, the blank rotates around the spindle axis; the position of the roller is controlled synchronously in both radial and axial directions during one revolution of the mandrel. When the mandrel rotates from 0\u00b0 to 180\u00b0, the roller moves from the top to the open end of the product in axial Parameters Values Pass angle \u03b8 (\u00b0) 60 direction, and the roller moves back axially when the mandrel rotates from 180\u00b0 to 0\u00b0 as shown in Fig. 3. Figure 4 shows the forming process using the translational pass set in asymmetric spinning. First, the product is formed from a sheet blank to shape (a) with an inclination angle \u03c6 by achieving different depths of the wall on two sides of the product. Next, the shape of the product is changed to (b) by keeping \u03c6 constant. Finally, \u03c6 is reducing to zero by achieving shape (c). The product is formed without a flange; all the material is pushed on the mandrel in the end. The experiments on cylindrical cups using the translational pass set in asymmetric spinning were conducted with the pa- rameters given in Table 1. For comparison, cylindrical cups using the same pass set in conventional spinning were formed. Figure 5 shows an example of the formed products using the translational pass set in asymmetric spinning. The thickness on the upper and lower sides of the wall (tU and tL as shown in Fig. 4) was measured axially. Figure 6 shows the axial distribution of wall thickness ratios (tU/t0 and tL/t0) with different pass pitches p = 2, 4, 6, and 8 mm. The dotted lines are distributions of thickness ratio (tC/t0) in conventional spinning (when \u03c6 = 0\u00b0). The shape of tU/t0 is shifted leftwards and tL/t0 is shifted rightwards when compared with tC/t0. Due to the inclination angle, material tends to flow towards the lower side of the wall at the beginning of the process which causes the different thickness distributions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000439_s40684-019-00177-3-Figure17-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000439_s40684-019-00177-3-Figure17-1.png",
        "caption": "Fig. 17 FE simulation of the rhombus core",
        "texts": [
            " The effects of the side layer on the elastic stiffness and strength are also calculated, with results shown in Fig.\u00a016a, b, respectively. When the raster angle is over 50\u00b0, the change of the raster angle does not affect the elastic stiffness and maximum force because the core deforms in rotation mode, as discussed above. Change of the side layer has a strong effect on the elastic stiffness and maximum force, as shown in Figs.\u00a016a, b. For further discussion, stress and plastic strain distributions are presented in Fig.\u00a017. ABAQUS software was employed for this finite element method (FEM), and a quarter model of the specimen, having a 2\u00a0T layer and 20% infill density, was modeled. Details of the FEM model are summarized in Table\u00a07. The FEM model is presented in Fig.\u00a017a, and Fig.\u00a017b, c shows 1 3 the results of the simulation at 2.1\u00a0mm of grip movement of the tensile test. As shown in Fig.\u00a017b, the side part has higher stress distribution than the core part. This result also supports the analytical solution and the experimental results, which showed that the side part plays an important role in resisting external forces. Figure\u00a017c shows plastic strain concentration on the joints in the core part. Finally, a study of the hexagonal core was conducted. Even though the analytical model was formulated based on the rhombus core, the physical concept (the hinge effect) can be applied to other specimens having different core structures. If the core part has joints, the hexagonal core 1 3 structures also will have the hinge effect. In addition, the side part also has a strong effect on the mechanical properties of the 3D printed specimen"
        ],
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    },
    {
        "image_filename": "designv11_63_0001197_s0263574720000533-Figure12-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001197_s0263574720000533-Figure12-1.png",
        "caption": "Fig. 12. The illustration of the ZMP and the support polygon for the snake robot.",
        "texts": [
            " px = Mgx + pzP\u0307x \u2212 L\u0307y Mg + P\u0307z (41) https://doi.org/10.1017/S0263574720000533 Downloaded from https://www.cambridge.org/core. University of Exeter, on 03 Jul 2020 at 02:30:12, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. py = Mgy + pzP\u0307y + L\u0307x Mg + P\u0307z (42) Then, we have pz = 0 considering that the snake robot locomotes on the ground. If the projection of the ZMP on the ground, (px, py) lies inside the support polygon, the snake robot is dynamically stable.24 The illustration is shown in Fig. 12. 3.1.1. The workspace of the predefined spiral curve method. Zhang et al. proposed a predefined spiral curve method for the head-raising of the snake robot.4 Then, a shape-fitting method based on the 3-DOF joint is proposed.4 And the roll, pitch, and yaw angles of the joint can be solved. But for the snake robot with single-DOF joints, it is not suitable. Therefore, we will compare the backbone curve based on the B\u00e9zier curve with that of a predefined spiral curve. And the modified discretization method is also applied to the predefined spiral curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000896_j.promfg.2020.04.265-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000896_j.promfg.2020.04.265-Figure2-1.png",
        "caption": "Fig. 2. Measurement process of the digital spotting image; (a) photogrammetry; (b) structured projection light scan; (c) color mesh [10].",
        "texts": [
            " In this respect, the digitalization of real spotting images and their evaluation represents a quantitative criterion by evaluating the content of the active surface during forming operations. The criterion is defined by an area content comparison between the target-bearing zone, from forming simulation, and the current active surfaces, from the tryout. For this approach, optical metrology from the GOM GmbH is used. By combining a photogrammetry measurement with a structured-light projection scan, a digital and three-dimensional spotting image is generated [8,9]. Color values are assigned to the individual STL-elements of the geometry scan via the beam paths of the SLR-camera. Fig. 2 shows the measurement process and a digital spotting image as its result. Due to the processing of the generated color meshes by a segmentation algorithm, the active surfaces can be extracted during die tryout. The color segmentation separates active areas from the rest of the mesh and analyzes them according to their surface contents. The aim is to compare the resulting surface contents with the given reference surfaces from the forming simulation. This comparison enables the evaluation of the current die performance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001689_1754337120961609-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001689_1754337120961609-Figure1-1.png",
        "caption": "Figure 1. A PIV image of a smooth ball with a wire trip. The vector background is vorticity making it possible to visualize the wake.",
        "texts": [
            " Conventional pitches, a category into which we will place any fastball (4-seam, 2-seam, sinker, cutter, and split finger), the curveball, changeup and slider, have a trajectory that is influenced by gravity and by the Magnus force.2,3 Knuckleballs, which do not rotate sufficiently fast to generate any appreciable Magnus force, are distinct. In this paper, we will discuss a newly discovered effect that, like Magnus force, stems from pressure differences due to a tilted wake. However, unlike Magnus force, this effect does not depend on the rotation of the ball. Rather, seams near the back of a baseball influence the flow symmetry over the ball. The Particle Image Velocimetry (PIV) dataset (described below) shown in Figure 1 is of a smooth, non-spinning ball with a wire \u2018\u2018trip\u2019\u2019 on its front to ensure the flow is turbulent everywhere. The pressure on a moving ball is highest where the air meets the front of the ball, which is called the \u2018\u2018stagnation point\u2019\u2019 (Pstag) in fluid dynamics. From Pstag the pressure decreases until the minimum pressure (Pmin) occurs where the surface of the ball is tangent to the flow. Separation of the boundary layer will occur on bluff bodies such as baseballs and other spheres leading to the formation of a wake"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002981_s12206-021-0627-8-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002981_s12206-021-0627-8-Figure11-1.png",
        "caption": "Fig. 11. Curvature center displacement under pure tilting load.",
        "texts": [
            " (16) It is defined here that when the outer raceway moves to the bearing center, the displacement is positive (squeezing the steel ball), and when the outer raceway moves away from the bearing center, the displacement is negative (out of contact). Therefore, the terms cosy i\u03b4 \u03c6 and sinz i\u03b4 \u03c6 have an opposite sign in Eqs. (15) and (16). When the ring is displaced under tiling load, assume that the tilt angle of the ring is z\u03b8 , as shown in Fig. 10. The outer raceway displacement zxi\u03b8\u03b4 and zyi\u03b8\u03b4 can be expressed as ( ) cos 2 2 1 cos 2 sin 2 mi m i mi zyi z mi zxi z d d d d \u03b8 \u03b8 \u03c6 \u03b4 \u03b8 \u03b4 \u03b8 \u23a7 = \u00d7\u23aa \u23aa \u23aa = \u2212\u23a8 \u23aa \u23aa =\u23aa \u23a9 . (17) The relationship of the raceway curvature center is shown in Fig. 11. So, the normal contact deformation ni\u03b4 between the steel ball and the raceway is, ' 0 0ni A A\u03b4 = \u2212 '2 ' 2 1 2 0d d A= + \u2212 . (18) ( ) ( ) 2 2 1 2 0zyi zxid d A\u03b8 \u03b8\u03b4 \u03b4= + + + \u2212 The relationship between the tilt displacement. z\u03b8 of the inner ring and the normal contact deformation ni\u03b4 can be established by Eqs. (17) and (18). Note that the absolute value of zyi\u03b8\u03b4 is taken in Eq. (18), because when the ring tilts, the raceway\u2019s y-direction displacement at each azimuth position all move toward the direction where the steel ball is squeezed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002948_j.jmapro.2021.05.040-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002948_j.jmapro.2021.05.040-Figure2-1.png",
        "caption": "Fig. 2. Channel arrangement on AM powder bed with build details.",
        "texts": [
            " CCC were fabricated using DMLS under the L-PBF category by using a laser melting system (Ren AM250, Renishaw, United Kingdom). Gas atomized AlSi10Mg powders with a mean diameter in the range of 25\u2013100 \u03bcm were used to fabricate the internal channels. The alloy used (AlSi10Mg) has high corrosion resistance. The presence of Silicon in the metal makes the alloy harder and more durable than pure aluminium due to the formation of Mg2Si precipitate. The build process is carried out in an argon environment to prevent oxidation. The layer thickness is kept constant at 40 \u03bcm. Fig. 2 illustrates the build details of the internal channels. The channels were built in the X-Y plane at 90\u25e6 build orientation with 5 mm thick support structures at the base. Fig. 3 shows the geometry of the CCC. The outer dimensions of the channels were kept constant at 9 \u00d7 9 mm. Three sets of internal channels with a cross-section (D) of 5 \u00d7 5, 3 \u00d7 3, and 1 \u00d7 1 mm were surface finished. Two sets of channel length (L) = 20 and 50 mm were considered for surface finishing. A new HCAF apparatus is designed and developed to surface finish the internal surfaces of channels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000934_s42835-020-00429-2-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000934_s42835-020-00429-2-Figure6-1.png",
        "caption": "Fig. 6 2D cross-section of the analysis model and its prototype",
        "texts": [
            " Motor parameters such as the resistance (1)Bpk = \u2211N\u22121 n=0 Bp(n)e j 2 kn N (2) m = ( Bp1 \u00d7 2 ) \u00d7 2 r P \u00d7 Lstk (3)Back EMF = \u221a 2 fkwN m Table 1 FEA results according to magnetizing current Magnetizing current (Arms) Fund. magnitude of air-gap flux density (T) Flux/pole (mWb) xm (Ohm) Fig. 4 Distribution of the air-gap flux density by stator and rotor winding 1 3 and reactance of each phase can be estimated from the stator geometry, rotor geometry, and winding details. Once these parameters are assigned to the corresponding impedance of the equivalent circuit, the stator and rotor currents that meet the desired speed or load requirements can be calculated. Figure\u00a06 shows a cross-sectional view of the analysis model for the two-dimensional (2D) FEA and a prototype of the model. This model has high magnetic saturation in the core because it is used for automotive applications that require high power within a limited size. It has 4 poles and 36 slots with distributed windings in the stator. The detailed specifications of the model are shown in Table\u00a02. The end ring of the rotor of an induction motor should be considered when performing the 2D FEA. Therefore, the analysis is performed by reflecting the end ring part in the conductivity of the secondary conductor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000679_edpc48408.2019.9012063-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000679_edpc48408.2019.9012063-Figure3-1.png",
        "caption": "Fig. 3 Combined rolling and bending process (adapted from [5])",
        "texts": [
            " Because of this, the same problems as in case of using rectangular wire material will occur here, but with the prior rolling, it is possible to manufacture different shapes for each winding if required. This is fundamentally important for being able to produce trapezoidal slots with high filling factors. To avoid the drawbacks of edgewise bending, a combined bending and rolling process is also conceivable. Therefore, [5] proposes to wind the wire around a rotating mandrel which represents the interior of the resulting winding body. At the same time the wire gets transferred to the desired shape by using a roll with defined cross-section (see Fig. 3). Through the superposition of the rotary motion and the feed motion of the mandrel, the wire is wound helically to the desired winding body. A forming roll rotating around its central longitudinal axis is provided for simultaneous crosssectional change. The rotating forming roll is arranged in such a way that at the point where the wire is bent around the mandrel, it exerts a forming force on the wire that changes the cross-section. However, every rolling process causes an elongation of the deformed material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001301_0142331220940427-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001301_0142331220940427-Figure2-1.png",
        "caption": "Figure 2. Force diagram of the satellite system.",
        "texts": [
            " The three-axis magnetorquer was completely designed and manufactured by Nanjing University of Science and Technology. Both the actuators meet the standards of CubeSat. A 3U CubeSat with these two actuators was assembled. Attitude dynamics of a satellite with moving masses The satellite was modeled as a rigid body with n moving masses and a three-axis magnetorquer. The following two Cartesian coordinate systems are here defined: an inertial coordinate system OI XI YI ZI and a body-fixed coordinate system OBXBYBZB. Figure 2 illustrates the pertinent variables required to derive the equations of motion. The resultant external force applied to the system is F, and Fb is the resultant external force applied to the rigid body (without moving masses). The force Fi is the net force acting on mass i. The force G is the gravity of the system and Fe is the environmental force applied to the system except G. The vector R locates OB in the OI XI YI ZI system, the vector pi locates the mass mi in the OBXBYBZB system, and the vector r locates any particle in the OBXBYBZB system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001363_s12206-020-0734-y-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001363_s12206-020-0734-y-Figure6-1.png",
        "caption": "Fig. 6. (a) Top-view vector profiles of computed features; (b) side-view flow vectors; (c) heat transfer rate to the ambient environment.",
        "texts": [
            " It should be emphasized that it is extremely difficult to experimentally measure the local temperatures of major rotor components during rotation, and the rotor magnets may have a temperature close to that of the airgap region if the airgap is thinner than the boundary layer, as will be discussed in more detail below. Various design factors of protrusion-shaped flow inducers on the rotor enclosure were considered, as shown in Fig. 2(b). The major design variables were the inclined angle (\u03b3), curvature (\u03ba), and height of the protrusion (H); the values of these variables were varied to determine the local convective cooling effect on the rotor magnets owing to the internal air circulation, as explained in Table 2. Fig. 6 depicts the predicted flow fields inside the entire motor domain, which sheds light on the internal operation of the electric motor. More specifically, Fig. 6(a) features the computed top-view vector profiles and Fig. 6(b) compares the side-view flow vectors for the cases with (left-half domain) and without (right-half domain) the protrusions, respectively. For a motor design without the protrusion, air exhibits relatively stronger radial vectors, owing to the dominant centrifugal force, which is attributed to the axial motor rotation, whereas there appears to be only negligible vertical flow motion through the ventilation holes in the rotor. In contrast, a motor design with the protrusion shows relatively weaker axial flow vectors, because part of the air is circulating through the ventilation holes (i.e., decrease in the axial momentum), in which there exists rather intensive vertical flow motion (i.e., increase in the vertical momentum), owing to the protrusion flow inducer. Fig. 6(c) shows the computed local heat transfer rates from the motor to the ambient environment, for the two representative cases with and without the protrusion, respectively, and under the same operating condition. For the motor design with the protrusion, the heat transfer rates from the motor frame, end bracket, and shaft to the ambient environment are higher, compared with the conventional protrusion-free design, because of the elevated temperature of the air that circulates through the ventilation holes of the rotor",
            "7 W), and was relatively higher compared with the thermal effect of the protrusion curvature on the radial heat transfer, shown in Fig. 8(b). It should be noted that the proposed flow inducers can increase the mass flow rate into the rotor ventilation holes such that the rotor magnet temperatures can be controlled under a designated temperature by the enhanced convective heat transfer rates. However, it should be noted that the protrusion results in relatively stronger axial flow disturbance (i.e., flow resistance) as shown in Fig. 6(a), which may require more input power to operate electric motors under the same operating conditions. In the motor design, one of the major design factors is the clearance between the stator and rotor, i.e., the airgap. Most of the heat generated in electrical motors moves in the radial direction, toward either the rotor frame or rotor shaft. When the Ohmic resistance dominantly dissipates the irreversible heat, the maximal temperature develops at the stator coils, and a part of the generated heat moves from the coil to the rotor magnet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002465_tasc.2021.3063644-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002465_tasc.2021.3063644-Figure6-1.png",
        "caption": "Fig. 6. The construction of the ring-shaped Cooling Dewar.",
        "texts": [
            " In the proposed RPS-HTSFS machine, the copper is chosen to armature windings and the HTS wires of the ring-shaped excitation coil are BSCCO-2223, whose operating temperature and current are 77 K and 120 A, respectively. The critical bend radius 30 mm is a limiting factor for the BSCCO-2223 HTS tape, but the use of ring-shaped coil structure can resolve this problem well. It\u2019s of necessity for the HTS coil to design the simple and high-efficient refrigeration system. Thanks to the configuration of the radial partitioned stator and single ring-shaped HTS coil, a special double-layer ring-shaped Dewar can be conceived to cooling the HTS wires. As shown in Fig. 6, the liquid nitrogen flows in from the lower pipe and out from the upper pipe, so that the single HTS-excitation coil placed in inner Dewar can be completely immersed in the cryogenic liquid nitrogen to keep the operation temperature. By using the poly-tetra-fluoroethylene (PTFE) support frame, the inner Dewar is fixed inside the outer Dewar. What\u2019s more, the vacuum between the inner and outer Dewar can guarantee a perfect thermal isolation. Fig. 7 shows the no-load back EMF waveforms at 300 rpm when the HTS-excitation current changes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002049_j.microc.2020.105902-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002049_j.microc.2020.105902-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of FI-CL system. PMT: photomultiplier tube; BPCL: ultra weak luminescence analyzer.",
        "texts": [
            "15 mL of H3PO4 and 20 mL of water were placed in a beaker and stirred evenly. Then the solution was transferred to a 50 mL Teflon-lined stainless steel autoclave and heated at 180 \u25e6C for 6 h in the blast drying oven. After the reaction, the reactor cooled to room temperature by naturally. The resulting dark yellow and transparent solution was subjected to dialysis in order to obtain the PClCDs. The synthesis method of P-doped CDs (PCDs) and Cl-doped CDs (ClCDs) was the same as that of PClCDs. The configuration of the FI-CL system in this operation was shown in Fig. 1. Two peristaltic pumps and a six-way injection valve were used to transport all the solution. The solutions of PClCDs and EY were pumped into the flow cell by a peristaltic pump 1 at the rate of 2.3 mL min\u2212 1. The Fe2+ solution, carrier H2O2 and standard or sample solution of Met were pumped by a peristaltic pump 2 at the rate of 2.5 mL min\u2212 1. The above solutions transported by two pumps were converged through a Y-shaped tube and then injected the flow CL reaction cell with the generation of CL signals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000786_s40430-020-2253-2-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000786_s40430-020-2253-2-Figure3-1.png",
        "caption": "Fig. 3 Meridian section structure of the tire Fig. 4 Finite element model of tire and drum test rig",
        "texts": [
            " In order to simulate the rolling process of the tire and ensure the consistency of the numerical simulation analysis and test processes, the rigid body is analyzed to define the drum. The contact boundary condition is defined between the tread and the drum [20]. The simplification of material partitioning is based mainly on two conditions: the size of the structural stress gradient or the deformation gradient and the degree of mechanical property difference of the adjacent regions. The meridian section structure of the tire is shown in Fig.\u00a03. In the simulation of the belt and ply, orthogonal anisotropic materials are used (using the embedded rebar unit) [21]. At the same time, the ply cords and the first and second belt cords are divided into finer grids to focus on the analysis. In the three-dimensional model, various rubber materials are represented by isotropic incompressible C3D8RH elements. The hyperelastic model is used to describe the stress\u2013strain relationship of the rubber material [22]. The number of carcass rubber units in the finite element model is 10,721, and the number of belt and cord layer units is 10,102"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000679_edpc48408.2019.9012063-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000679_edpc48408.2019.9012063-Figure5-1.png",
        "caption": "Fig. 5 Forming tool for sequential forming",
        "texts": [
            " To avoid the previously described drawbacks a new forming process was developed at Fraunhofer IWU [6], [7]. In this process, first the wire is sequentially pressed into the necessary shape. Afterwards it is wound with conventional bending machines and finally calibrated in the head or the slot area of the coil. In this paper the first forming step will be described in detail. III. TOOL FOR SEQUENTIAL FORMING Producing a coil by sequential forming requires a forming tool, which is easily adaptable to the proposed geometries. The realized forming tool at Fraunhofer IWU is shown in Fig. 5. It produces sequences directly from a raw wire coil and consists of two wire straightening units (horizontal and vertical), a NCfeed with lifting function, two horizontal axes to adjust the width of the sequence and two pressing jaws whose height is adjusted with two distances. The jaws have a flat area, which represents the length of the slot and a transition zone. Depending on the shape of this transition zone a lengthening of the wire occurs while pressing. To avoid buckling the feed has to release the wire while pressing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002568_1.4042545-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002568_1.4042545-Figure1-1.png",
        "caption": "Fig. 1. Reference configuration of the 4-bar linkage (from [9]).",
        "texts": [
            " Application of the properties of the derivatives of a determinant gives rise to the general expression (26) for the differentials of the minors, where the truncated multi-index ak = (a1,a2, . . . ,ak) \u2208 N k is defined by removing the last n\u2212 k indexes from a, and ni := |{a j|a j = i}| is the number of times a differential of degree i occurs. The differentials of S\u03b1\u03b2i are determined by (25). The first example is the planar 4-bar linkage with four revolute joints, i.e. V \u2282T 4. In the (singular) reference configuration q0 = 0, shown in fig. 1, the n= 4 joint screw coordinates are Y1 = Y3 = (0,0,1,0,L,0)T ,Y2 = (0,0,1,0,2L,0)T Y4 = (0,0,1,0,0,0)T . Analysis of V : The kinematic tangent cone to V was shown to be [9] CK q0 V = V(x1 + x3,x2,x4)\u222aV(x1 + x2,x3 + x2,x4 \u2212 x2). That is, the linkage can perform smooth 1-DOF motions through q0, and two motion branches intersect at that point. It is known that the configuration q0 is a bifurcation of V , and CK q0 V is in fact the union of the two tangent spaces to the manifolds intersecting at q0",
            "\u201d Differential and Combinatorial Topology, Princeton University Press (1965): pp. 205-244 [25] de Jong, Jan J, Mu\u0308ller, Andreas and Herder, Just L: Higher-order Taylor approximation of finite motions of mechanisms.\u201d Robotica (2018): published online. DOI 10.1017/S0263574718000462 Acc ep te d Man us cr ip t N ot C op ye di te d Downloaded From: https://mechanismsrobotics.asmedigitalcollection.asme.org on 02/01/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use List of Figure Captions Figure 1 Reference configuration of the 4-bar linkage (from [9]). Figure 2 a) Planar linkage, constructed by connecting two Watt linkages, in its reference configuration q0 (from [10]). b) Topological graph and fundamental cycles. Figure 3 2nd-order approximation V 2 q0 of the c-space of the linkage in fig. 2: a) x1-x5 section, b) x1-x2 section. Figure 4 Two nearby configurations the linkage can reach from the cusp singulaity q0, where a) corresponds to the red part, and b) to the blue part of the curve in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure46.1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure46.1-1.png",
        "caption": "Fig. 46.1 HSS die and punch set",
        "texts": [
            " The rpm of the lathe is fixed at 240 rpm and rotated clockwise for 30 min and then anticlockwise for another 30 min for dispersion of copper powder in aluminium powder. After 1 h of rotation, the container is taken out of the chuck, steel balls are separated and mixed powder is weighted again. An error of 0.010 g weight before and after mixing is observed at that time. The mixed powder of aluminium with 10% of copper is filled in the hollow cylinder of a high speed steel (HSS) die as shown in Fig. 46.1. The inner diameter of the die is 16 mm and outer diameter of the die is 50 mm. The height of die and punch is of 55 mm and 61 mm, respectively. The pin seats on the base to stop displacement of die and base. Then the die is placed on the base. After filling the die with powder, compaction is done with the help of punch. Seat of die which is made of aluminium is used to extract the cold compacted sample from die. Depending upon the compaction pressure and volumeof powder, cylindrical shape sample of height 16 mm and diameter 16 mm is prepared"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000124_012164-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000124_012164-Figure5-1.png",
        "caption": "Figure 5. Displacement",
        "texts": [],
        "surrounding_texts": [
            "Tegangan ekivalen yang digunakan metode Von-Mises. Berikut ini ilustrasi hasil analisis equivalent stressTegangan ekivalen maksimum terjadi di bagian las rangka bagian depan sebesar 16.32 MPa, kemudian tegangan ekivalen minimum sebesar 0.01 MPa."
        ]
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure57.4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure57.4-1.png",
        "caption": "Fig. 57.4 Meshing of the component, a Using a seed size of 5 mm throughout, b Finer meshing detailing at the cushion mounting regions",
        "texts": [
            " The topology of the component was checked for smooth meshing of the component using four different criteria: free edges, solid shells, shell faces, and wire edges. As a standard guideline, smaller geometries (edges shorter than 0.1 mm and faces with areas smaller than 1 mm2) were merged with adjacent faces. Geometry edit tools were used to rectify/eliminate the imprecise geometry in the component. Figure 57.3b presents the model after clean-up using edit tools. The isotropic model approach was chosen to assign the resultant laminate properties such as Young\u2019s modulus and Poisson\u2019s ratio obtained from the experiments. A global seed size of 5 mm (Fig. 57.4a) was used for the rider seating region of the component, and seed size of 3 mm was used for rubber cushion mounting holes and the guide for rubber cushion (Fig. 57.4b). This seed size was varied depending on the thickness of the component. In order to arrive at the boundary conditions, a seat load impression test was carried out using blue print method on existing seat base under the load conditions of: nopassenger, solo passenger, and twin passenger configurations. The boundary conditions were applied at the initial step in the CAD model (Fig. 57.5). The seat base bottom rubber cushion mounting and the rear rubber grommet mounting region were constrained along the vertical direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002912_s12555-020-0049-x-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002912_s12555-020-0049-x-Figure2-1.png",
        "caption": "Fig. 2. RPR configuration of STS mechanism.",
        "texts": [
            " DYNAMIC MODELLING OF SIT-STAND MECHANISM A dynamic model is necessary for the simulation and analysis of a system, because it is the first step toward control implementation. The mathematical model of the STS mechanism describes its motion between the sitting and standing positions. The ARP CAD presented in Fig. 1 includes two parts: an STS mechanism to support the shift of a patient from a sitting to a standing position and vice versa and the mobile platform controlled by a joystick. Khan and Kausar discussed the kinematic modelling of ARP [30]. The robot mechanism is considered to be the revolute\u2013prismatic\u2013revolute (RPR) mechanism shown in Fig. 2. The Lagrange approach implements the equation of motion of the system [31]. This dynamic equation can be expressed as (1): M(\u03b8)\u03b8\u0308 +C(\u03b8 , \u03b8\u0307)\u03b8\u0307 +N(\u03b8 , \u03b8\u0307) = \u03c4, (1) where \u03c4 , M, C and N are actuator torque vector, mass matrix, Coriolis matrix and gravity term respectively. The equations of motion for ARP [25] are given as (2) to (4). \u03c41 = (m1l2 1 + Izz1 +m2d2 2 + Iyy2 +m3l2 c3 + Ixx3S2 3 + Iyy3C2 3)\u03b8\u03081 +2m2d2d\u03072\u03b8\u03071\u2212m2d2\u03b8 2 1 +(m1l1 +m2d2)gC1 +m3glc3(S1S3\u2212C1C3), (2) \u03c42 = (m2 +m3)d\u03082\u2212m2d2\u03b8\u0307 2 1 +m2gS1, (3) \u03c43 = (m3l2 c3 + Izz3)\u03b8\u03083\u2212 (Ixx3S3C3\u2212 Iyy3S3C3)\u03b8\u0307 2 1 +m3gle3(C1C3\u2212S1S3), (4) where d2 and \u03c4i are the displacements of the linear actuator and torque at i-th joint respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001541_ccc50068.2020.9189251-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001541_ccc50068.2020.9189251-Figure1-1.png",
        "caption": "Figure 1: The self-driving forklift hardware system",
        "texts": [
            " 5540 Authorized licensed use limited to: Cornell University Library. Downloaded on September 16,2020 at 01:09:05 UTC from IEEE Xplore. Restrictions apply. all the discrepancy of the model from the forklift is lumped as total disturbance to be observed [15] and rejection in realtime. Experimental validation results will be provided and performance metrics in terms of tracking error and settling time in different test scenarios will be presented. An in-house developed self-driving forklift is developed based on the Linde E20 forklift as shown in Figure 1. The coordinates are measured by the ultra-high bandwidth (UWB) system suitable for indoor applications. Note that to reduce the hardware cost, only one UWB module is mounted on the forklift. The heading angle is obtained through the attitude heading reference system (AHRS) sensor. The steering and vehicle velocity are controlled by an electrical motor. Radar system is mounted for obstacle avoidance. The driving brain is an in-house developed control unit based on embedded control system. The flow chart for lateral control is shown in Figure 2",
            " The following observations are clear 1) For the case with = 15 rad/s, with the a1 varying from -5 to -50, both Pm and Gm decreases. 2) Increasing to 30 rad/s, Pm and Gm are barely affected by the variation of a1. 3) Comparing the results in 1) and 2), elevated helps improving the robustness to plant dynamic variations. 4) In practical applications, however is limited by facts such as low sampling rate. Overall, frequency domain analysis in this section showed satisfactory robustness against plant dynamic variations. The robustness improves with increased . Experimental results based on the forklift shown in Figure 1 were discussed in this section. In this test, only the inner loop controller was enabled, where the target heading angle is set as 90 degree. It was seen from Figure 12 that the actual heading angle initiates from 140 degree and converges to 90 1.0 degree after 7s. The variation of steering angle was within 40 degree. The 5545 Authorized licensed use limited to: Cornell University Library. Downloaded on September 16,2020 at 01:09:05 UTC from IEEE Xplore. Restrictions apply. smooth tracking process confirmed the effectiveness of the inner-loop controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002163_s40430-020-02796-3-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002163_s40430-020-02796-3-Figure2-1.png",
        "caption": "Fig. 2 Test bench",
        "texts": [
            " Finite Difference Method is used to divide m \u00d7 n grids uniformly along circumferential direction and axial direction of lubricating film, for total of (m + 1) \u00d7 (n + 1) grid nodes, m = 400 and n = 40. Discrete equation Solving Eq.\u00a0(22) using successive over-relaxation (SOR) Iteration Method: where \u03b2 is super relaxation factor, 1 < \u03b2 < 2; k is number of pressure iterations. When pressure is iterated to a certain number of times, program stops iterating when formula (25) is satisfied \u03b6 is pressure convergence accuracy, is 1.0 \u00d7 10\u201310. Diagram of simulation calculation process is shown in Fig.\u00a01. The picture of test setup is shown in Fig.\u00a02, bearing ring block is shown in Fig.\u00a03, and installation diagram of bearing ring block is shown in Fig.\u00a04. Friction torque of bearing ring block can be obtained by torque equipment as shown in Fig.\u00a02. During tests, load P (specific pressure, Pa) is set to 0.2\u00a0MPa by weight Fblock (unit: N). (22) B1i,jpi+1,j + B2i,jpi\u22121,j + B3i,jpi,j+1 + B4i,jpi,j\u22121 \u2212 B5i,jpi+1,j = B6i,j (23) \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 B1i,j = \ufffd h 3 \ufffd i+0.5,j B2i,j = \ufffd h 3 \ufffd i\u22120.5,j B3i,j = \ufffd d l \ufffd2\ufffd \u0394 \u0394 \ufffd2 \ufffd h 3 \ufffd i,j+0.5 B4i,j = \ufffd d l \ufffd2\ufffd \u0394 \u0394 \ufffd2 \ufffd h 3 \ufffd i,j\u22120.5 B5i,j = B1i,j + B2i,j + B3i,j + B4i,j B6i,j = 3\u0394 \ufffd\ufffd h \ufffd i+0.5,j \u2212 \ufffd h \ufffd i\u22120.5,j \ufffd (24) p k i,j = ( B1i,jp k\u22121 i+1,j + B2i,jp k\u22121 i\u22121,j + B3i,jp k\u22121 i,j+1 + B4i,jp k\u22121 i,j\u22121 \u2212 B6i,j B5i,j ) + p k\u22121 i,j (25) \u2211n j=2 \u2211m i=2 \ufffd\ufffd\ufffdpki,j \u2212 pk\u22121 i,j \ufffd\ufffd\ufffd\u2211n j=2 \u2211m i=2 \ufffd\ufffd\ufffdpki,j \ufffd\ufffd\ufffd \u2264 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:56 1 3 Page 5 of 10 56 where P is specific pressure of bearing ring block, Pa; S is area of bearing ring block, m2; R2 is outer diameter of bearing ring block, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002546_s12541-021-00485-2-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002546_s12541-021-00485-2-Figure2-1.png",
        "caption": "Fig. 2 Contact pressuredistribution of the micro contact",
        "texts": [
            " 1 Elastic contact between asperities rough surface deformed asperities rigid surface mean plane 1 3 In this section, a tangential damping model of bolted joint based on the physics-based coefficient is presented. The contact area of the micro-contact can be divided into stick region and slip one. For the bolted joints, the energy dissipation of contact surfaces mainly comes from the slip region of the micro-contact. The dissipation can be represented as the tangential damping mathematically. The pressure distribution of the micro-contact under a constant normal load and oscillating tangential load can be shown as Fig.\u00a02. The study [24] has shown that the radius of the stick region reduces with the increase of tangential force when the normal force is invariant. The ratio of the stick-regime radius c to real contact radius r is defined as follows where Qx,P are the tangential and normal forces respectively applied on the micro contact, is the static friction coefficient, which can be defined as the ratio of maximum tangential load Qmax to the normal load Pmax as [21] where r is the equivalent shear stress. (12)A\u2217 r = Ar Aa = D (2\u2212D)\u22152a \ufffdD\u22152 l 2(2 \u2212 D)Aa ( a \ufffd(2\u2212D)\u22152 l + a\ufffd(2\u2212D)\u22152 c ) (13)c r = ( 1 \u2212 Qx P )1\u22153 (14) = Qmax Pmax = rAr HAr The tangential deformation of a single asperity under the constant normal load and oscillating tangential load can be given as where G\u2217 = (( 2 \u2212 v1 ) \u2215G1 + ( 2 \u2212 v2 ) \u2215G2 )\u22121 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure7-1.png",
        "caption": "Fig. 7. Temperature distribution for slot current density J=16.4 A(rms)/mm2",
        "texts": [],
        "surrounding_texts": [
            "The LPTN is used to provide a quick estimate of temperatures at different locations in the motor for different slot current densities. These were then compared against the temperatures obtained from CFD analysis to verify the validity of the LPTN model. Temperature rise in different portions of the machine using both CFD and lumped parameter network are shown in Table IV. It is visible that CFD results match closely with that of the LPTN model and the continuous output power for this 120 kW slotless machine using the proposed WELC method is 85 kW when the temperature rise is limited to 80 \u25e6C. This also corresponds to a current density of 23.3 A/mm2. This validates that the proposed WELC method is effective to further increase the maximum allowable continuous current density compared to the state-of-the-art cooling methods, which makes it possible to improve the power density as well. To analyze the effectiveness of thermal conductivity of the winding support material, the hotspot temperature of the winding is studied for different thermal conductivities ranging from 3 W/m-K (similar to Zirconia) to 18 W/m-K (similar to Silicon Nitride), as shown in Fig. 11. It is found that the continuous current density can be increased using materials with higher thermal conductivity. It is also found that, the benefits of higher thermal conductivity saturates at higher values."
        ]
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure3-1.png",
        "caption": "Figure 3. Meshing conditions of paralleled gear pair.",
        "texts": [
            " The coordinates of tooth face for gear 1 and gear 2 at this moment can be expressed as follows, respectively Xo g1 =T1X 1 g1 \u00f03\u00de Xo g2 =T2X 2 g2 \u00f04\u00de where T1 is the rotation matrix from O1 x1y1z1 to Oo xoyozo and T2 is the transformation matrix from O2 x2y2z2 to Oo xoyozo, which can be described as follows, respectively T1 = cosuc1 sinuc1 0 0 sinuc1 cosuc1 0 0 0 0 1 0 0 0 0 1 2 664 3 775 \u00f05\u00de T2 = 1 0 0 0 0 cosp sinp 0 0 sinp cosp 0 0 0 0 1 2 664 3 775 1 0 0 0 0 1 0 a 0 0 1 0 0 0 0 1 2 664 3 775 cosuc2 sinuc2 0 0 sinuc2 cosuc2 0 0 0 0 1 0 0 0 0 1 2 664 3 775 \u00f06\u00de Since it is a line contact form for paralleled beveloid gear pair, each slice has one contact point if we divide the gear into ns slices, ns ! \u2018 pieces. For a certain moment uc1 and a certain piece si of gear 1, we can calculate the contact point location by considering the following conditions based on the gear mesh theory: 1. Contact points of both gears should have the same position coordinate in the same coordinate system; 2. Normal vectors at the contact point of both gears should have the same director. The conditions mentioned above shown in Figure 3 can be expressed as Xo g1 =Xo g2 jnog13nog2j=0 \u00f07\u00de where nog1 and nog2 denote the normal vectors of Xo g1 and Xo g2, respectively. We still follow the slicing method above (when ns ! 0, each slice can be regarded as normal evolute cylindric gear), then the potential energy in slice si of jth meshing gear pair of gear 1 and gear 2 can be expressed as Ui 1j and Ui 2j, respectively. The expression of the above assumption is Ut = min Pnp 1 Pns 1 Ui 1j +Ui 2j \u00f08\u00de where Ut is the total potential energy in gear pair at moment t, np is the number of multiple teeth meshing at the same time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002979_s10514-021-09996-3-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002979_s10514-021-09996-3-Figure4-1.png",
        "caption": "Fig. 4 Palm contact transition model, PC1, represented as a set of projection vector from the approximated shoulder point. Given pt , we first identify approximated shoulder point with a fixed transform from pt , and then project palm contact from the shoulder points using projection vectors in PC1",
        "texts": [
            " For palm contact, we first define an approximate torso pose pt as pt = [ xl f +xr f 2 yl f +yr f 2 zl f +zr f 2 0 0 \u03b8l f +\u03b8r f 2 ]T (1) [ xl f , yl f , zl f ] and [ xr f , yr f , zr f ] are the left and right foot positions, respectively, and \u03b8l f and \u03b8r f are the rotations of each foot about the z axis. We can then derive approximated shoulder points based on the approximated torso pose. The palm contacts are projected from each approximated shoulder point to the environment with a predefined set of projection vector, as shown in Fig. 4. Since our PFS is a search-based planner, a cost is required for each action. For the foot action, we define the cost function as g f = dt + ws , where dt is the distance the approximated torso travels in this action and ws is a fixed cost of taking a step. For the palm action, the cost is gp = dp +ws , where dp is the distance the palm travels in this action. Each state is feasible if there exists a collision-free and statically-balanced inverse kinematics solution for the specified end-effector poses",
            " For each \u03b3 , if m is a motion mode which requires Algorithm 1: Compute Ground Truth Label for Traversability Input : pt (Torso pose), v (Torso translation),m (Motion mode),E (Environment, specified as a set of surfaces),FC1 (Foot contact transition model) ,PC1 (Palm contact transition model); \u2190 GetFeasibleFootstepCombinations (pt ,E,FC1); + \u2190 { }; for \u03b3 in do if ContactSequenceExists (\u03b3, v,m,E,FC1,PC1) then + \u2190 {\u03b3 } \u222a +; end end return | +|; palm contacts, the plannerwill start from a set of initial states, whose foot locations are \u03b3 , with all combinations of possible palm contacts given by the palm transition model PC1, as shown in Fig. 4. The above process is line 5 in Algorithm 1. In this work, we quantify the traversability as the number of useful footstep combinations, denoted | +|, which serves as an indicator for how difficult it is for the planner to find a contact transition sequence in environment E moving from torso pose pt with v torso translation using motion mode m. The process computing | +| is summarized in Algorithm 1. However, to compute the traversability in an environment E would require running a PFS planner for every combination of (pt , v,m), which is clearly too time-consuming",
            " Finally we sum the scores for each \u03b1 to obtain the footstep score S f . The palm contact direction also affects the robot\u2019s ability to move in a certain direction. For example, if there exist only palm contacts in the direction opposite to the direction of motion, the planner may not be able to find a valid contact transition sequence. Therefore, we add an additional feature: the contact clearance feature of palm contacts around pt . We first find all palm contact projection onto the environment with a palm contact transitionmodelPC1, as shown in Fig. 4. Each projection then returns a nearest contact point c on one of the surfaces. To represent the direction of the palm contacts relative to pt , we divide the palm scores into the four quadrants of the torso frame: Sp[i] for the i th quadrant. This process corresponds to Lines 10 to 14 in Algorithm 2. With [ S f , Sp[1], Sp[2], Sp[3], Sp[4] ] , we define the contact clearance feature vector S (pt , v,m,E) as: S (pt , v,m,E) = \u23a7 \u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 [ S f ] , m = feet only[ S f , Sp[1], Sp[2] ] , m = feet and left palm[ S f , Sp[3], Sp[4] ] , m = feet and right palm[ S f , Sp[1], Sp[2], Sp[3], Sp[4] ] , m = all end-effectors (13) To train the estimator for each motion mode, we generate multiple environment with randomly tilt surfaces, and collect ground truth data for the difficulty in planning using the PFS approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002216_s12541-020-00457-y-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002216_s12541-020-00457-y-Figure2-1.png",
        "caption": "Fig. 2 Pneumatic (left) [22] and electric (right) [23] clamping device",
        "texts": [
            " Pennock and Israr [18] designed the link lengths of a six-bar linkage mechanism wherein an adjustable pivot is employed. Huang et\u00a0al. [19] defined a design parameter for a five-bar double-toggle clamping mechanism and optimized the link lengths of the mechanism using a genetic algorithm. Goulet and Li [20] synthesized four-bar linkage using non-convex optimization. Gawande and Bhojane [21] optimized toggle clamp mechanism used in CNC bending machines by analytical and numerical way. The most common type of clamping device is the DESTACO [22] shown in Fig.\u00a02 (left), which uses pneumatic cylinders. It uses air pressure to clamp and unclamp the finger. Although pneumatic clamps are widely used, they have two principal limitations. First, workspace of the manipulator is limited due to complex pneumatic tubing for air lines. Second, a large tank to generate air pressure is needed. The electric clamp is being developed as an alternative to overcome the limitations of the pneumatic clamp. The electric clamp uses only electric wire and does not need air tubes. This eliminates the use of air tanks and frees the workspace, making the whole system simpler (Fig.\u00a02 (right)). In addition to the actuation method of the clamping device, the type of linkage mechanism used enables toggling features. The four-bar linkage is the most common linkage mechanism. Figure\u00a03 [24] shows the internal structure of the common pneumatic clamp with a four-bar linkage mechanism. The pneumatic cylinder applies force to link 1 with prismatic motion and rotates links 2 and 3. Link 3 rotates around the output joint and generates an output torque and clamping force. When the device is fully clamped, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000115_j.mechmachtheory.2019.103647-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000115_j.mechmachtheory.2019.103647-Figure3-1.png",
        "caption": "Fig. 3. Motorcycle coordinates.",
        "texts": [
            " The rotation of the front wheel is not considered due to the computational cost of solving the tire-road interaction twice. Thus, the convergence of the simulation has been increased as only a pure traction scenario is considered. Neither aerodynamic forces nor rolling resistance have been considered in order to be able to analyse the phenomena under study independently. A total of 39 coordinates are used to model the motorcycle ( Table 1 ). 30 natural coordinates describe the movement of the 15 points shown in Fig. 3 . The remaining 9 coordinates are grouped into two types: lengths associated with the four damper-spring systems considered and five angles. Three angles are associated with the transmission and the other two with the pitch of the chassis and swingarm. Table 1 Coordinates. Coordinate Description Coordinate Description x 1 Swingarm axis x 11 Wheel sprocket tangential point y 1 y 11 x 2 Engine axis x 12 Upper point fork suspension y 2 y 12 x 3 Upper damper pin x 13 Front wheel axis y 3 y 13 x 4 Upper triple clamp x 14 Rear tire contact point y 4 y 14 x 5 Lower triple clamp x 15 Front tire contact point y 5 y 15 x 6 Rear wheel axis rs Rear suspension length y 6 fs Front suspension length x 7 Lower damper pin rt Rear tire radii y 7 ft Front tire radii x 8 Engine reference point \u03b8 Chassis angle y 8 \u03b1 Engine sprocket angle x 9 Rear wheel reference point \u03b2 Wheel sprocket angle y 9 \u03b3 Upper chain angle + \u03c0 /2 x 10 Engine sprocket tangential point \u03c6 Swingarm angle y 10 Table 2 Inertial properties of each body",
            " Therefore, the previously described angular movement has to be taken into account to properly model the whole system. This statement is of great importance in the present work because it shows that, as angular accelerations appear in the drivetrain, power transmission from the engine to the wheel is strongly influenced by the design of the final drive. Three different road profiles are simulated in this work, namely: 1. Sinusoidal profile 2. Gaussian profile (bumpy) 3. Flat profile The coordinates of points 14 and 15 ( Fig. 3 ) are given by equations (32) and (33) according to the type of road: sinusoidal, Gaussian and flat. y 14 \u2212 H y sin ( 2 \u03c0 \u03bb x 14 ) = 0 (32a) y 14 \u2212 H y exp ( \u2212 (x 14 \u2212 \u03bc) 2 2 \u03c3 2 ) = 0 (32b) y 14 = 0 (32c) y 15 \u2212 H y sin ( 2 \u03c0 \u03bb x 15 ) = 0 (33a) y 15 \u2212 H y exp ( \u2212 (x 15 \u2212 \u03bc) 2 2 \u03c3 2 ) = 0 (33b) y 15 = 0 (33c) The forces that govern the dynamics of the system are grouped into three distinct blocks: front and rear suspensions, vertical forces on the front and rear tires and longitudinal forces on the rear tire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001809_ecce44975.2020.9235728-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001809_ecce44975.2020.9235728-Figure1-1.png",
        "caption": "Fig. 1. Discussed SEIG system.",
        "texts": [
            " Section II presents the mathematical model of induction generator which considers the magnetic saturation effects, variable stator/rotor leakage split ratio and unbalanced condition. Section III demonstrate the emulation system and parameter 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 1794 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 30,2021 at 06:51:47 UTC from IEEE Xplore. Restrictions apply. measurements of the induction machine. Section VI has shown test results of emulation, simulation, and experiment to verify the emulator under unbalanced conditions. Finally, Section V concludes the work. Fig. 1 shows the discussed SEIG system. The renewable energy source which in this case, the wind turbine is the prime-mover and input power to the induction generator. The excitation capacitors are connected to the induction generator in parallel with the loads, helping machine to build up voltage. The generated power is supplied to the loads, either balanced or unbalanced. This paper focuses on the system performance under unbalanced load conditions. System models are established for the induction generator, which consider the magnetic saturation effects, excitation capacitors, and load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure11-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000507_b978-0-12-817463-0.00012-5-Figure11-1.png",
        "caption": "Fig. 11 The center of mass of links 1 and 2.",
        "texts": [],
        "surrounding_texts": [
            "Knee and hip rehabilitation were independently modeled and controlled in PHYSIOTHERABOT. The dynamic analysis and control method is explained here. Knee joint dynamic analysis: The general dynamic model of a robot manipulator was given in Eq. (11). The knee joint is a single DOF and can be modeled like a pendulum driven by the axis of rotation as shown in 417Impedance control applications in therapeutic exercise robots xr zr xs2 zs2 xs1 zs1 xr zr q2 q1 418 Erhan Akdogan and Mehmet Emin Aktan 419Impedance control applications in therapeutic exercise robots The coordinate of the knee sensor connected to the link 2 and the Jacobian matrix: The coordinates in the x- and z-axes and partial derivatives of the sensor are given in the following equation: x2 \u00bc l1 cos\u03b81 + r1 cos\u03b82! \u03b4x2\u00bc l1\u03b4\u03b81 sin\u03b81 r1\u03b4\u03b82 sin\u03b82 z2 \u00bc l1 sin\u03b81 r1 sin\u03b82! \u03b4z2\u00bc l1\u03b4\u03b81 cos\u03b81 r1\u03b4\u03b82 cos\u03b82 (38) \u03b4x2 \u03b4z2 \u00bc l1 sin\u03b81 r1 sin\u03b82 l1 cos\u03b81 r1 cos\u03b82 \u03b4\u03b81 \u03b4\u03b82 Jacobian matrix for link 2 is given in Eq. (39). The transpose, inverse, and derivative of the Jacobian are given in Eqs. (40)\u2013(42), respectively. J2 \u00bc l1 sin\u03b81 r1 sin\u03b82 l1 cos\u03b81 r1 cos\u03b82 (39) JT2 \u00bc l1 sin\u03b81 l1 cos\u03b81 r1 sin\u03b82 r1 cos\u03b82 (40) J 1 2 \u00bc r1 cos\u03b82 r1 sin\u03b82 l1 cos\u03b81 l1 sin\u03b81 1 l1r1 sin\u03b81 cos\u03b82 + l1r1 cos\u03b81 sin\u03b82 (41) _J 2 \u00bc l1\u03b4\u03b81 cos\u03b81 r1\u03b4\u03b82 cos\u03b82 l1\u03b4\u03b81 sin\u03b81 r1\u03b4\u03b82 sin\u03b82 (42) Dynamic equations: For calculating the joint torques of the robot manipulator, the dynamic equations of the system must be calculated. In cases where no external force is not applied, the Lagrange equation given in the following equation is used in the joint torque calculation: \u03c4\u00bc d dt \u03b4K \u03b4 _\u03b8 \u03b4K \u03b4\u03b8 + \u03b4P \u03b4\u03b8 (43) where P and K are the potential and kinetic energy, respectively. The coordinates of the robot manipulator parts: In order to calculate the kinetic and potential energies in the Lagrange equation, the positions and velocities of the robot manipulator and lowerlimb parts (Figs. 11 and 13) must be found. These equations are given in the following: Part 11: x11\u00bc lg11 cos\u03b81! _x11 \u00bc lg11 _\u03b81 sin\u03b81 z11\u00bc lg11 sin\u03b81! _z11\u00bc lg11 _\u03b81 cos\u03b81 (44) 420 Erhan Akdogan and Mehmet Emin Aktan Hip apparatus (ha) part: xha \u00bc 0:25cos\u03b81! _xha \u00bc 0:25 _\u03b81 sin\u03b81 zha \u00bc 0:25sin\u03b81! _zha\u00bc 0:25 _\u03b81 cos\u03b81 (45) Thigh part: xthigh \u00bc lgthigh cos\u03b81! _xthigh \u00bc 0:144lhuman sin\u03b81 z11\u00bc lgthigh sin\u03b81 ! _zthigh \u00bc 0:144lhuman cos\u03b81 lgthigh \u00bc 0:144lhuman (46) Part 22: x22\u00bc\u00f0l1 b+ lg22\u00decos\u03b81 + _l21 cos\u03b82 _x22\u00bc \u00f0l1 b+0:2\u00de _\u03b81 sin\u03b81 0:2 _\u03b82 sin\u03b82 z22\u00bc\u00f0l1 b+ lg22\u00de sin\u00f02\u03c0 \u03b81\u00de _l21 sin\u03b82 _z22\u00bc \u00f0l1 b+0:2\u00de _\u03b81 cos\u03b81 0:2 _\u03b82 cos\u03b82 _l21\u00bc 0:2 \u00bdm (47) 421Impedance control applications in therapeutic exercise robots Shank part: xshank \u00bc l1 cos\u03b81 + lgshank cos\u03b82! _xshank \u00bc l1 _\u03b81 sin\u03b81 lgshank _\u03b82 sin\u03b82 zshank \u00bc l1 sin\u03b81 lgshank sin\u03b82 ! _Z shank\u00bc l1 _\u03b81 cos\u03b81 lgshank _\u03b82 cos\u03b82 lgshank \u00bc 0:144lhuman (48) Foot part: xfoot \u00bc l1 cos\u03b81 + l2 cos\u03b82 + lgfoot cos \u03c0 2 \u03b82 ! _xfoot \u00bc l1 _\u03b81 sin\u03b81 l23 _\u03b82 sin\u03b82 lgfoot _\u03b82 sin \u03c0 2 \u03b82 zfoot \u00bc l1 sin\u03b81 lg2 sin\u03b82 + lgfoot sin \u03c0 2 \u03b82 ! _zfoot \u00bc l1 _\u03b81 cos\u03b81 lg23 _\u03b82 cos\u03b82 lgfoot _\u03b82 sin \u03c0 2 \u03b82 lgfoot\u00bc 0:429lfoot (49) where lhuman is height of patient, lgthigh, lgfoot, and lgshank are mass of centers of thigh, foot, and shank, respectively. The kinetic energy of the robot manipulator parts: In order to calculate the kinetic energy used in the Lagrange equation, the kinetic energies of the robot manipulator components are calculated separately and given in the following: Part 11: K11\u00bc 1 2 m11l 2 g11 _\u03b8 2 1 + 1 2 IG11 _\u03b8 2 1\u00bc 0:122 _\u03b8 2 1 (50) Hip apparatus(ha) part: Kha \u00bc 1 2 Ioha _\u03b8 2 1\u00bc 0:132 _\u03b8 2 1 (51) Thigh part: Kthigh\u00bc 1 2 Ithigh _\u03b8 2 1 \u00bc 10 7\u00bd\u00f021,186BW 222,796\u00de+20:736mthighl 2 human (52) where BW represents body weight and m is mass of part. Part 21: K21\u00bc 1 2 m21\u00f0 _x221 + _z221\u00de2 + 1 2 IG21 _\u03b8 2 2\u00bc 1:625\u00f0 _x221 + _z221\u00de+0:04 _\u03b8 2 2 (53) 422 Erhan Akdogan and Mehmet Emin Aktan Part 22: K22\u00bc 1 2 m22\u00f0 _x222 + _z222\u00de2 + 1 2 IG22 _\u03b8 2 2\u00bc 0:375\u00f0 _x222 + _z222\u00de+2:83 _\u03b8 2 1 (54) Part 23: K23 \u00bc 1 2 m23\u00f0 _x223 + _z223\u00de2 + 1 2 IG23 _\u03b8 2 2 \u00bc 0:9\u00f0 _x223 + _z223\u00de+4:6587 10 4 _\u03b8 2 2 (55) Shank part: Kshank\u00bc 1 2 mshank\u00f0 _x2shank + _z2shank\u00de2 + 1 2 IGshank _\u03b8 2 2 \u00bc 10 7\u00f05341BW +44,749\u00de (56) Foot part: Kfoot \u00bc 1 2 mfoot\u00f0 _x2foot + _z2foot\u00de2 + 1 2 IGfoot _\u03b8 2 1 \u00bc 10 7\u00f0355BW +7296\u00de (57) The potential energy of the robot manipulator parts: In order to calculate the potential energy used in the Lagrange equation, the potential energies of the robot manipulator components are calculated separately and given the following: Part 11: P11\u00bcm11gz11\u00bc 5:75sin\u03b81 (58) Part hip apparatus: Pha \u00bcmhagzha \u00bc 10:423sin\u03b81 (59) Part thigh: Pthigh \u00bcmthighgzthigh \u00bc 1:412\u00f0mthighlhuman\u00de sin\u03b81 (60) Part 21: P21 \u00bcm21gz21\u00bc 31:88\u00f00:68sin\u03b81 + lg21 sin\u03b82\u00de (61) Part 22: P22 \u00bcm22gz22 \u00bc 7:3575\u00bd\u00f00:68 b+ lg22\u00de sin\u00f02\u03c0 \u03b81\u00de 0:2sin\u03b82 (62) Part 23: P23\u00bcm23gz23\u00bc 1:7658\u00f00:68sin\u03b81 + 0:09sin\u03b82\u00de (63) 423Impedance control applications in therapeutic exercise robots Shank part: Pshank \u00bcmshankgzshank \u00bc mshank9:81\u00f00:68sin\u03b81 + 0:144lhuman sin\u03b82\u00de (64) Foot part: Pfoot \u00bcmfootgzfoot \u00bc mfoot9:81 0:68sin\u03b81 + 0:5sin\u03b82 lfoot0:429sin \u03c0 2 \u03b82 (65) The contribution of the parts to joint torques: The contributions of the parts to the joint torques calculated using the Lagrange equation: For Joint 1: \u03c4111 \u00bc 0:244\u20ac\u03b81 5:75cos\u03b81 (66) \u03c41ha \u00bc 0:265\u20ac\u03b81 10:243cos\u03b81 (67) \u03c41thigh \u00bc Iothigh\u20ac\u03b81 1:412mthighlhuman cos\u03b81 (68) \u03c4121 \u00bc 3:25l21 \u20ac\u03b81 + 3:25l1lg1\u20ac\u03b82 cos\u00f0\u03b81 \u03b82\u00de +3:25l1lg1 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de +3:25l1lg1 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 21:67cos\u03b81 (69) \u03c4122 \u00bc\u00f00:75\u00f0l1 b+0:2\u00de2 + 5:66\u00de\u20ac\u03b81 + 0:15\u20ac\u03b82 cos\u00f0\u03b81 \u03b82\u00de +0:15 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 0:15 _\u03b81 _\u03b82\u00f0l1 b+0:2\u00decos\u00f0\u03b82 \u03b81\u00de +7:3575\u00f00:6 b+ lg22\u00decos\u03b81 (70) \u03c4123 \u00bc 1:8l21 \u20ac\u03b81 + 1:8l1lg23\u20ac\u03b82 cos\u00f0\u03b81 \u03b82\u00de+1:8l1lg23 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de +1:8l1lg23 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 1:2cos\u03b81 (71) \u03c41shank \u00bcmshankl 2 1 \u20ac\u03b81 + lgshank \u20ac\u03b82 cos\u00f0\u03b81 \u03b82\u00de+ l1lgshank _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de +mshankl1lgshank _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de mshank6:6708cos\u03b81 (72) \u03c41foot \u00bc l21 \u20ac\u03b81 + Ifoot \u20ac\u03b81 +mfootl1 \u20ac\u03b82\u00f0Bc\u03b81 A sin\u03b81\u00de+mfoot\u00bd2l1l2 _\u03b822 cos\u03b81 sin\u03b82 + 2 _\u03b8 2 2l1lgfoot sin\u03b82 cos\u03b81 2 _\u03b82l1 _\u03b81B sin\u03b81 + 2 _\u03b8 2 2 sin\u03b81l1lgfoot sin\u03b82 + 2 _\u03b8 2 2l1l2 sin\u03b82 sin\u03b81 2 _\u03b82 _\u03b81l1Ac\u03b81 mfoot _\u03b82 _\u03b82\u00f0l2 sin\u03b82 sin\u03b81 + lgfoot sin\u03b82 sin\u03b81\u00de+2 _\u03b82 _\u03b81\u00f0 sin\u03b81lgfoot cos\u03b82 + sin\u03b81l2 sin\u03b82\u00de 6:67cos\u03b81 (73) 424 Erhan Akdogan and Mehmet Emin Aktan A \u00bc lgfoot sin \u03c0 2 \u03b82 l2 sin\u03b82 (74) B \u00bc l2 cos\u03b82 lgfoot sin \u03c0 2 \u03b82 (75) For Joint 2: \u03c4211\u00bc \u03c42ha\u00bc \u03c42thigh \u00bc 0 (76) \u03c4221 \u00bc 3:25l1lg1\u20ac\u03b81 cos\u00f0\u03b81 \u03b82\u00de+3:25lg21\u20ac\u03b82 + 0:08\u20ac\u03b82 + 3:25l1lg1 _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 3:25 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 31:88lg21 cos\u03b82 (77) \u03c4222 \u00bc 0:15\u00f0l1 b+0:2\u00de\u20ac\u03b81 cos\u00f0\u03b81 \u03b82\u00de+0:03\u20ac\u03b82 + 0:15 _\u03b81\u00f0l1 b+0:2\u00de\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 0:15 _\u03b81 _\u03b82\u00f0l1 b+0:2\u00de sin\u00f0\u03b81 \u03b82\u00de 1:4715cos\u03b82 (78) \u03c4223 \u00bc 1:8lg23l1 \u20ac\u03b81 cos\u00f0\u03b81 \u03b82\u00de+1:8lg23 \u20ac\u03b82 + 4:6587 10 4\u20ac\u03b82 + 1:8l1lg23 _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 1:8l1lg23 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 0:16cos\u03b82 (79) \u03c42shank\u00bcmshankl1lgshank \u20ac\u03b81 cos\u00f0\u03b81 \u03b82\u00de+ \u20ac\u03b82\u00f0mshankl2gshank + Igshank\u00de +mshankl1lgshank _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de mshankl1lgshank _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de mshank1:41264cos\u03b82 (80) \u03c42foot \u00bcmfootl1 \u20ac\u03b81 cos\u03b81B+mfoot \u20ac\u03b82B 2 +mfootl1 _\u03b81 Bcos\u03b81 _\u03b81B sin\u03b81 mfoot _\u03b81 _\u03b82l1 l2 sin\u03b82 cos\u03b81 lgfoot cos \u03c0 2 \u03b82 cos\u03b81 lgfoot cos \u03c0 2 \u03b82 sin\u03b81 + l2 cos\u03b82 sin\u03b81 mfoot9:81 0:5cos\u03b82 lfoot0:429sin \u03c0 2 \u03b82 (81) Joint torques: The joint torques can be calculated by using the sum of the contributions of the parts to the joint torque and using the following equations: \u03c41 \u00bc \u03c4111 + \u03c41ha + \u03c41thigh + \u03c4121 + \u03c4122 + \u03c4123 + \u03c41shank + \u03c41foot (82) \u03c42 \u00bc \u03c4211 + \u03c42ha + \u03c42thigh + \u03c4221 + \u03c4222 + \u03c4223 + \u03c42shank + \u03c42foot (83) Robot dynamic equation for hip exercises: In the case of external force acting, the dynamic equation of a robot manipulator is as in Eq. (11). Coriolis, gravity, and other effects and inertial matrix M can be calculated using Eqs. (82), (83), (11). The inertia matrix in Eq. (11) and all the information in the hN vector are given in the equations between Eqs. (66) and (81). Accordingly, the M(\u03b8) matrix 2\u00de 2\u00de 425Impedance control applications in therapeutic exercise robots M \u00bc M11 M12 M21 M22 (84) M11 \u00bc 0:244+ 0:265+ Iothigh + 3:25l2 1 + \u00f00:75\u00f0l1 b+0:2\u00de2 + 5:66\u00de+1:8l2 1 + l2 1 + Ifoot (85) M12 \u00bc 3:25l1lg1 cos\u00f0\u03b81 \u03b82\u00de+0:15cos\u00f0\u03b81 \u03b82\u00de+1:8l1lg23 cos\u00f0\u03b81 \u03b82\u00de +lgshank cos\u00f0\u03b81 \u03b82\u00de+mfootl1\u00f0Bcos\u03b81 A sin\u03b81\u00de (86) M21 \u00bc 3:25l1lg1 cos\u00f0\u03b81 \u03b82\u00de+0:15cos\u00f0l1 b+0:2\u00decos\u00f0\u03b81 \u03b82\u00de+1:8lg23l1 cos\u00f0\u03b81 \u03b82\u00de +mshankl1lgshank cos\u00f0\u03b81 \u03b82\u00de+mfootBl1 cos\u03b81 (87) M22 \u00bc 3:25lg21 + 0:11 + 1:8lg23 + 4:6587 10 4 +mshank l 2 gshank + Igshank +mfootB 2 (88) hN\u00f0\u03b8, _\u03b8\u00de vector: hN \u00f0\u03b8, _\u03b8\u00de\u00bc hN1\u00f0\u03b8, _\u03b8\u00de hN2\u00f0\u03b8, _\u03b8\u00de (89) hN1\u00f0\u03b8, _\u03b8\u00de\u00bc 5:75cos\u03b81 10:243cos\u03b81 1:412mthighlhuman cos\u03b81 + 3:25l1lg1 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b8 +3:25l1lg1 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 21:67cos\u03b81 + 0:15 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 0:15\u00f0l1 b+0:2\u00de _\u03b82 _\u03b81 cos\u00f0\u03b82 \u03b81\u00de+7:3575\u00f00:6 b+ lg22\u00decos\u03b81 + 1:8l1lg23 _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de+1:8l1lg23 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 1:2cos\u03b81 + l1lgshank _\u03b82\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de+mshankl1lgshank _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de mshank6:6708cos\u03b81 +mfoot\u00bd2 _\u03b82c1l1l2 sin\u03b82 + 2 _\u03b822l1lfoot\u03b81 sin\u03b82 cos\u03b81 2 _\u03b82l1 _\u03b81B sin\u03b81 +2 _\u03b82 sin\u03b81l1lgfoot sin\u03b82 + 2 _\u03b822l1l2 sin\u03b81 sin\u03b82 2 _\u03b82 _\u03b81l1Acos\u03b81 mfoot _\u03b81 _\u03b82\u00f0l2 sin\u03b82 sin\u03b81 + 2 _\u03b81 _\u03b82\u00f0 sin\u03b81lgfoot + sin\u03b81l2\u03b82\u00de\u00de 6:67cos\u03b81 hN2\u00f0\u03b8, _\u03b8\u00de\u00bc 3:25l1lg1 _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 3:25 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 31:88lg21 cos\u03b82 + 0:15 _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de\u00f0l1 b+0:2\u00de 0:15\u00f0l1 b+0:2\u00de\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 1:4715cos\u03b82 + 1:8l1lg23 _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de 1:8l1lg23 _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b82\u00de 0:16cos\u03b82 +mshank l1lgshank _\u03b81\u00f0 _\u03b82 _\u03b81\u00de sin\u00f0\u03b81 \u03b82\u00de mshank l1lgshank _\u03b81 _\u03b82 sin\u00f0\u03b81 \u03b8 mshank1:41264cos\u03b82 +mfoot _\u03b81l1\u00f0 _B cos\u03b81 _\u03b81B sin\u03b81\u00de+mfoot _\u03b81 _\u03b82l1\u00f0 l2 sin\u03b82 cos\u03b81 lgfoot cos \u03c0 2 \u03b82 cos\u03b81 lgfoot cos \u03c0 2 \u03b82 sin\u03b81 + l2 cos\u03b82 sin\u03b81\u00de mfoot9:81 0:5cos\u03b82 lfoot sin \u03c0 2 \u03b82 0:429 Jacobian vector Jx: Jx \u00bc Jx1 Jx2 (90) Ta Ex P A Is Is Is R 426 Erhan Akdogan and Mehmet Emin Aktan The equations related to the Jacobian matrix are given in Eqs. (37), (39), (40). The external forces Fext acting on PHYSIOTHERABOT are given in Eqs. (33), (34)."
        ]
    },
    {
        "image_filename": "designv11_63_0002332_iccss52145.2020.9336861-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002332_iccss52145.2020.9336861-Figure3-1.png",
        "caption": "Fig. 3. Spiral hinge connection between back side and bottom surfaces",
        "texts": [],
        "surrounding_texts": [
            "In response to the above-mentioned problems, this paper designs a variable-angle side-phase anti-bird cage. The new side-phase anti-bird cage is aimed at specific types of transmission line towers, taking into account the possible differences between the design drawings of the transmission tower and the reality, and the angle change is used to achieve the attachment of the device to the transmission line tower, ensuring that the protected area of the device will not be allowed birds enter the nest. The shape of the device is a triangular prism composed of five grid surfaces, in which the shapes of the bottom surface and the top surface are isosceles trapezoids, and the sizes of the two are the same, which is convenient for mass production during processing. The back side is a stretchable structure composed of two parts. The sides are two triangular mesh surfaces of the same size. The grid size on all sides of the device is 10mm*10mm. This size can ensure that even the smallest birds cannot enter the nest protected by the device in the area where the device is placed."
        ]
    },
    {
        "image_filename": "designv11_63_0003203_s42835-021-00864-9-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0003203_s42835-021-00864-9-Figure9-1.png",
        "caption": "Fig. 9 The spatial magnetic field intensity distribution of the receiving coil (a) surface intersecting structure, (b) edge intersecting structure",
        "texts": [
            " Finite element method (FEM) is a numerical technique for solving the approximate solutions of boundary value problems of partial differential equations. It can make full use of the convenience provided by high-speed computer to solve the problems quickly.In this paper, FEA is used to calculate the surface intersecting structure and the edge intersecting structure. The transmitting coil is excited with 10 A current. The spatial magnetic field intensity distribution of the receiving coil is shown in Fig.\u00a09. It can be observed from the results that the magnetic field distribution at the intersection areas of adjacent independent coils in the two structures is different. In addition, in order to more clearly compare thedifference of the spatial magnetic field distribution between the surface intersecting structure and the edge intersecting, coils, as shown in the Fig.\u00a09. The distribution of magnetic field intensity on the section l of the two structures is shown in the Fig.\u00a010. As can be seen from the figure, the change the intercept line l is taken at the center of the four receiving trend of flux density under the two rectangle, and the magnetic flux density is the largest at the position directly above the receiving coil C1 and C2. Comparing the two structures, the magnetic flux density value of the rectangle surrounded by the receiving coil C1, C2, C3, and C4 is basically the same, while the magnetic flux density value of the surface intersecting structure is significantly greater than that of the edge intersecting structure at the position directly above the receiving coil C1 and C2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001299_s00202-020-01062-y-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001299_s00202-020-01062-y-Figure1-1.png",
        "caption": "Fig. 1 Components of a canned motor",
        "texts": [
            ":(0123456789) Keywords Canned motor\u00a0\u00b7 Electrical equivalent circuit\u00a0\u00b7 Thermal model\u00a0\u00b7 Insulation systems\u00a0\u00b7 Nuclear applications Canned motor pumps are hermetic pumps mainly used in nuclear and chemical areas for pumping radioactive, poisonous, and flammable fluids. The canned motor has a thin, cylindrical, and non-magnetic metallic pipe \u201cthe can\u201d inserted in air gap to prevent the pumped fluid from entering the stator and damaging windings. The pumped fluid is used to cool the motor and lubricate the hydrodynamic bearings as illustrated in Fig.\u00a016 (see \u201cAppendix\u201d) The rotor is also canned with a non-magnetic material to prevent corrosion as shown in Fig.\u00a01. The double-canned motor presents many advantages such as small starting current, environmental protection, extremely low noise and vibrations, low life cycle cost, and high mechanical strength [1, 2]. However, it also has some limitations. In fact, \u201cthe air gap\u201d (filled with the pumped fluid) of canned motor is wide compared to usual induction motors and the losses are higher due to the generation of eddy currents in the cans and the mechanical losses due to the friction of the fluid on rotative parts (Fig.\u00a01). In addition, the stator lifetime is given by the maximum temperature of the windings which strongly depends on the operating conditions of the motor. The new standards on electrical motors efficiency, the increase in the service life of nuclear power plants, as well as the safety growth of installations after the Fukushima disaster, are the reasons why canned motor pumps must have better efficiency and lifetime while guaranteeing safety. Therefore, the design of such canned motors must be improved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001470_2050-7038.12588-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001470_2050-7038.12588-Figure6-1.png",
        "caption": "FIGURE 6 The physical structure of the magnet",
        "texts": [
            " In this case, the air gap between the stator and the magnet will take a minimum value at the center and a maximum value at the edges, and a varying air gap will be obtained.38 The design problem of cogging torque with offset amount can be formulated as follows: Minimize Tcog offset\u00f0 \u00de Subjected to : Prated\u2265P3kW \u00f07\u00de The initial value for offset parameter was determined as 0 mm, the upper limit as 50 mm and the sensitivity as 1 mm, and analyses were performed. The structure of the magnet when offset is 0 and 50 mm is presented in Figure 6. As the amount of offset increases, the amount of collapse at the edges of the magnet also increases. Figure 7 shows the change in cogging torque and generator power in relation to offset amount. With the value of pole arc offset, cogging torque and output power decrease. The change is cogging torque displayed a similar behavior to skew application. However, in this case, the reduction in the volume of the magnet resulted in a significant decrease in the output power. At the lowest cogging torque, the value of the output power decreased to as much as 2560 W"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001361_aim43001.2020.9158838-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001361_aim43001.2020.9158838-Figure1-1.png",
        "caption": "Fig. 1. The model of USV.",
        "texts": [
            " Thus, the overall effect of nonlinearities, modeling uncertainties and external disturbance is well handled. Finally, the stability and effectiveness 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 1298 Authorized licensed use limited to: University of Wollongong. Downloaded on August 10,2020 at 20:48:18 UTC from IEEE Xplore. Restrictions apply. of proposed control design is guaranteed with the lyaponov stability theory and comparative simulation, and the simulate results show the tracking priority of proposed control design. II. THE MODELING OF USV\u2019S OVERALL SYSTEM With the model of USV in Fig.1, the kinematic model and nonlinear dynamic model are derived as follows: \u03b7\u0307 = R (\u03c8)V (1) M0\u03b7\u0308+C0\u03b7\u0307+D0\u03b7\u0307 = R\u03c4 + dm + ds (2) where \u03b7 = [ x y \u03c8 ]T is the USV\u2019s position, where x, y represent the global coordination and \u03c8 represents the heading angle of the USV. V= [ u v r ]T is the velocity state of USV, where u, v and r represent the surge, sway velocities and yaw rotational velocity respec- tively. R (\u03c8)= \u23a1 \u23a3 cos\u03c8 \u2212 sin\u03c8 0 sin\u03c8 cos\u03c8 0 0 0 1 \u23a4 \u23a6 is the rotation ma- trix. M0 = RM \u2032 0R T , C0 = R ( C \u2032 0 \u2212M \u2032 0R T R\u0307 ) RT , D0 = RD\u2032 0R T , dm is modeling uncertainties and can be estimated by fuzzy logic system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure4-1.png",
        "caption": "Figure 4. Parameters needed in solving potential energy: (a) parameters of normal cylinder gear and (b) parameters of slice si of beveloid gear.",
        "texts": [
            " According to the research,24 the total potential energy of ordinary cylindrical gear teeth at a certain moment in the meshing process can be divided into Hertz contact potential energy Uh, bending potential energy Ub, shearing potential energy Us, axial compression potential energy Ua and tooth foundation deformation energy Uf, each part of potential energy can be expressed as Uh = F2 1 n2\u00f0 \u00de 2pEb \u00f09\u00de Ub = \u00d0d 0 Min 2 2EIx dx \u00f010\u00de Us = \u00d0d 0 1:2Ft 2 2GAx dx \u00f011\u00de Ua = \u00d0d 0 Fa 2 2EAx dx \u00f012\u00de Uf = F2 2bE cosact\u00f0 \u00de2 L uw Sfw 2 +M uw Sfw +P 1+Q tanact\u00f0 \u00de2 \" # \u00f013\u00de where F is the contact force between gear pair, Fa is the radial force, Ft is the tangential force, Min is the input torque of the meshing gear pair, act denotes the pressure angle at the contact point, b denotes the tooth width of the gear, d denotes the height of contact point, Ax denotes the cross-sectional area, Ix denotes the moment of inertia. Parameters uw and Sfw are shown in Figure 4, the L, M, P and Q are the coefficients listed in the literature.25n, G and E are Poisson\u2019s ratio, shear modulus and Young\u2019s modulus, respectively. Based on results from TCA, we define the contact point coordinates on slice si of jth meshing gear pair of gear 1 as Xi 1j considering the existence of multiple teeth participation in meshing at the same time. We notice that each equation of each part of potential energy can be expressed as Ux = cuxF 2 \u00f014\u00de where Ux is the x type potential energy, cux can be seen as a constant of each part of potential energy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001435_10.0001388-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001435_10.0001388-Figure5-1.png",
        "caption": "Fig. 5. A pictorial representation of the torques for a precessing football moving in the positive z direction with gravity turned off. Shown on the left is the start of the motion with a deviation of s\u0302 in the positive x direction. The torque due to the airflow is shown as a directed circular arc. This torque in the positive y direction results in the configuration shown on the right, in which s\u0302 has a deviation in the positive y direction. In this figure, the aerodynamic (yaw) torque, now in the negative x direction, is again shown as a directed circular arc.",
        "texts": [
            " The heart of the resolution of our football paradox lies in the response of the football to the aerodynamic torque caused by the non-zero angle of attack that develops after its launch, assumed for this discussion to be that of a perfect spiral pass. To illustrate this idea, we neglect gravity and assume that the football is moving in a straight line in the \u00fez direction. If both s\u0302 and L are in the z-direction, nothing interesting happens, and so we now tilt the front of the football slightly upward, i.e., in the positive x direction. (This is the equivalent of the trajectory turning downward under the action of gravity in the actual pass.) The airflow then imposes a torque in the \u00fey\u0302 direction as shown in Fig. 5(a). Since L is essentially in the same direction as s\u0302, the torque causes s\u0302 to develop a component along \u00fey\u0302; i.e., gyroscopic precession results in the ball yawing right. In this position, as shown in Fig. 5(b), the aerodynamic torque develops a component in the x direction, giving rise to a component of precession about y\u0302, and so the ball starts to tip downward. (In the case of an actual pass immediately after launch, this torque along x results in a slowing of the growth of the SX component (see Fig. 4)). Put simply, the gyroscopic precession resulting from the pitch- and yaw-induced aerodynamic torque, coupled with the co-linearity of L and s\u0302, causes dSY=dt / SX (Fig. 5(a)) and dSX=dt / SY (Fig. 5(b)), which means that the tip of S moves in a circle with an angular frequency xgyr: With this simple case in mind, we now consider what happens in a real pass after the ball has been launched. Again, we suppose that v\u0302, L, and s\u0302 are initially aligned. After the launch, however, v\u0302 begins to slowly rotate downward under the action of gravity. By \u201cslowly,\u201d we mean that the rate at which the velocity direction rotates is significantly smaller than the other relevant rotation rate, xgyr. (In our example, the rate of v\u0302 rotation is about 0",
            "19 s, in acceptably good agreement with the 0:17 s period for xwob \u00bc 38.0 rad/s. Interestingly, because of the initial pitch, the gyroscopic precessional radius now increases from the perfect-pass case of 0.15 to 0.22. Thus, the average deviation angle between s\u0302 and v\u0302 increases significantly. Importantly, however, the center of gyroscopic precession remains the same at a zero-pitch yaw of 9 ; the ball still tracks the line of trajectory closely. The path of S in Fig. 7 can be compared directly with the result shown in Fig. 5 of Rae\u2019s paper.1 That figure shows roughly 1.5 slow oscillations during the 3 s flight and roughly 20 high-frequency cycles. We note that our results agree with those of Rae in showing that the direction of the symmetry axis is vertically symmetric, i.e., symmetric about SX \u00bc 0, just as Rae\u2019s results are symmetric about a \u00bc 0. However, in our calculation, the yaw angle does veer into the \u2013y territory for a brief period between 1.7 s and 2.2 s into its flight. There is a significant difference between the pattern of high-frequency loops in Fig. 5 in the study by Rae and the pattern of cusps in Fig. 7. The difference can be ascribed to shifts in the phase of the pitch and yaw. It is interesting that the looping is almost absent early in Fig. 5 in the study by Rae, suggesting that it requires details such as the effect of air drag on the shape of the trajectory not included in our model; that effect might be significant only late in the trajectory. In analytically exploring the model of Eqs. (2) and (4), the actual path of a football under gravity is a complicating factor since the rate at which v\u0302 rotates is variable. We will, therefore, sacrifice some reality for simplicity by using a \u201ccircular\u201d model for the change of v\u0302, v\u0302z \u00bc cos Xt\u00f0 \u00de; v\u0302x \u00bc sin Xt\u00f0 \u00de; (10) where X is the angular velocity of the rotation of v\u0302 in the xz plane and is taken to be constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001530_0954406220957369-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001530_0954406220957369-Figure2-1.png",
        "caption": "Figure 2. Schematic of a flexible rotor-bearing system.",
        "texts": [
            " The expressions for force components in \u2018X\u2019 and \u2018Y\u2019 directions are written below: FX \u00bc Z L 0 Z 2p 0 pRcosh dhdz (20a) FY \u00bc Z L 0 Z 2p 0 pRsinhdhdz (20b) Resultant internal force (Fint) on the journal due to film pressure is evaluated using the following relation: Fint \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 X \u00fe F2 Y q (21) Shear force (Fs) is computed using the following relation:3 Fs \u00bc Z L 0 Z 2p 0 h 2R @p @h \u00fe g U h Rdhdz (22) Coefficient of friction (f) is calculated using follow- ing relation: f \u00bc Fs Fint (23) Dynamic characteristics of rotor-bearing system To study the dynamic characteristics of a rotorbearing system, it is important to know the stiffness and damping coefficients of the bearing. Thus, the lubricating film is modeled to have both direct and cross coupled stiffness and damping elements as illustrated in Figure 2. A finite perturbation approach is used to find the stiffness and damping coefficients. Figure 3 shows the concept of finite perturbation where the centre of the journal (Oj) is given a very small perturbation of displacement and velocity from its steady state position (Oj0). Perturbation amplitudes of 0.002Cr and 0.002Crxs are taken for displacement and velocity perturbations respectively in the present paper. The resulting bearing forces are evaluated, and stiffness and damping coefficients are obtained as: KXX \u00bc FX\u00f0X0 \u00fe DX;Y0; 0;0\u00de FX\u00f0X0 DX;Y0; 0;0\u00de\u00bd 2DX\u00f0 \u00de (24a) KXY \u00bc FX\u00f0X0;Y0 \u00fe DY; 0;0\u00de FX\u00f0X0;Y0 DY; 0;0\u00de\u00bd 2DY\u00f0 \u00de (24b) KYX \u00bc FY\u00f0X0 \u00fe DX;Y0; 0;0\u00de FY\u00f0X0 DX;Y0; 0;0\u00de\u00bd 2DX\u00f0 \u00de (24c) Equation (25) can be written in the normalized form as M X 00 \u00fe ks X X0\u00f0 \u00de \u00fe cs X 0 X 0 0 \u00bc 0 (26a) M Y 00 \u00fe ks Y Y0\u00f0 \u00de \u00fe cs Y 0 Y 0 0 \u00bc 0 (26b) 2 KXX X0 \u00fe 2 KXY Y0 \u00fe 2 CXX X 0 0 \u00fe 2 CXY Y 0 0 ks X X0\u00f0 \u00de cs X 0 X 0 0 \u00bc 0 (26c) 2 KYX X0 \u00fe 2 KYY Y0 \u00fe 2 CYX X 0 0 \u00fe 2 CYY Y 0 0 ks Y Y0\u00f0 \u00de cs Y 0 Y 0 0 \u00bc 0 (26d) where, M \u00bc MCrxs 2 W ; ks \u00bc ksCr W ; cs \u00bc csCrxs W ; Kij \u00bc KijCr W ; Cij \u00bc CijCrxs W ; X \u00bc X Cr ; Y \u00bc Y Cr ; X0 \u00bc X0 Cr ; Y0 \u00bc Y0 Cr ; X 0 \u00bc d X ds ; Y 0 \u00bc d Y ds ; X 0 0 \u00bc d X0 ds ; Y 0 0 \u00bc d Y0 ds ; X 00 \u00bc d2 X ds2 ; Y 00 \u00bc d2 Y ds2 ; X 00 0 \u00bc d2 X0 ds2 ; Y 00 0 \u00bc d2 Y0 ds2 ; s \u00bc xst Equations (26a) to (26d) is rewritten in the matrix for as below: M X 00f g \u00fe C X 0f g \u00feK Xf g \u00bc 0 (27) where, M, C, and K are the mass, damping, and stiffness matrices given by M \u00bc M 0 0 0 0 M 0 0 0 0 0 0 0 0 0 0 2 664 3 775; C \u00bc cs 0 cs 0 0 cs 0 cs cs 0 2 CXX \u00fe cs 2 CXY 0 cs 2 CYX 2 CYY \u00fe cs 2 664 3 775; K \u00bc ks 0 ks 0 0 ks 0 ks ks 0 2 KXX \u00fe ks 2 KXY 0 ks 2 KYX 2 KYY \u00fe ks 2 664 3 775 Ob Oj0 (X0 0,Y ) Oj (X +0 0+ X,Y Y\u0394 \u0394 ) e X Y For cases where the rotor is assumed to be rigid, the displacement of disc centre X;Y\u00f0 \u00de will be same as the displacement of journal centre X0;Y0\u00f0 \u00de, as the disc is supposed to be placed in the middle of the shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002895_14644207211019767-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002895_14644207211019767-Figure1-1.png",
        "caption": "Figure 1. 3D model of (a) the exterior face of the door; (b) the interior frame of the door; (c) the entire door.",
        "texts": [
            "19 Mathematical modeling of the door structure To build the car door, it was necessary to make two pieces, an exterior face and an interior frame. The first phase in the realization of the exterior face of the door consisted in the dimensional sampling of a car door, of the same type, made of metal, with which to then make a comparison in terms of behavior. The Appendix presents the technical drawings of the interior and exterior parts and the door assembly. These drawings were be used to make the threedimensional (3D) model of this landmark. Figure 1 shows the 3D model of the outer face of the door, the inner face, and the door assembly. The modeling was performed using the dimensions from the technical drawing presented in the Appendix (see Figures 16 to 18). The second phase in the process of making the outer face of the door was the construction of the mold in which the molding was done. The mold was made of polyester resin, reinforced with fiberglass (Figure 2). The outer face was made of two layers of carbon fiber fabric and a layer of polyester on the inside"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002736_s13177-021-00256-3-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002736_s13177-021-00256-3-Figure1-1.png",
        "caption": "Fig. 1 Final lateral distance (d) between the vehicle and the bottom of the parking spot",
        "texts": [
            " It should be noted that the vehicle must stop at tangent point to change the steering in this method. G\u00f3mez-Bravo et al. developed a solution to make the algorithm independent from initial position of vehicle in [14]. Vorobieva et al. proposed two newmethods to plan a path for autonomous vehicles to park in tiny spots with the same approach [15\u201317]. The earlier studies about autonomous parking of nonholonomic vehicles using geometric methods have a common constraint which is the limit of the minimum final lateral distance to the parking spot. This distance is illustrated as \u201cd\u201d in Fig. 1. The minimum value of \u201cd\u201d depends on the dimensions of the parking spot and the vehicle parameters. Especially for longer vehicles like busses, coaches, trucks etc., classical geometric approaches result with the high values of \u201cd\u201d. This makes these geometric methods inapplicable and unsuitable for long vehicles. One of the contributions of this paper is to examine the performances of the classical geometric approaches with respect to the mentioned lateral distance (d) for the first time in literature",
            " proposed 2 methods using similar manner in 2012 [15]. Detailed test results of this work can be found in [16]. All of the work including the control of the steering signals are summarized by Vorobieva et al. in 2015 [17]. These 2 geometric methods are called as \u201cpark in one maneuver\u201d, and \u201cpark in several reversed trials\u201d are the basis of our work and they are explained for completeness in the next subsections. Detailed information about these methods can be found in [15\u201317]. As it is seen in Fig. 1 A(xA, yA), J(xJ, yJ), B(xB, yB), D(xD, yD) and E(xE, yE) points are fixed to the vehicle where A, J, B, D points are the corners and E point is the center of the rear track of the vehicle. Additionally, G(xG, yG) and F(xF, yF) points are the lower-left and the upper-right corner of the parking spot. (See Fig. 1). Also, all the methods in this paper are presented for the scenario where the parking spot is on the right side. In order to park the vehicle to its left, same equations can be modified using symmetry. Since the parking maneuver is performed at low speeds, kinematic equations provide the motion model sufficiently accurate. Figure 2 shows the kinematic model where (x, y) denote the position of center of the rear axle and \u03c8, \u03b4 stands for the orientation angle and the steering angle of the vehicle respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002411_iros45743.2020.9340935-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002411_iros45743.2020.9340935-Figure2-1.png",
        "caption": "Fig. 2: Design of the BioARS (a) Module: inchworm robot. (b) CAD model of the driving mechanism of the inchworm robot. (c) Assembled robot: quadruped robot.",
        "texts": [],
        "surrounding_texts": [
            "An assembly robotic system includes a swarm of robots as modules, which bind with each other through connectors. The assembly process of the system is generally task-oriented and the assembled robot has better adaptability and more functions compared with a single module. Therefore, the design and fabrication of an assembly robotic system is always focused on the module, the connector, and the assembled robot. In the BioARS, the inchworm robot is the basic module. A micro-controller and battery are attached to the inchworm robots, making the entire system untethered. Magnets are used to connect the inchworms. We use a \u201cshoulder-to-shoulder\u201d connecting method to prevent failure due to large bending moment during assembly and walking. The assembled robot is a quadruped which consists of eight inchworm robots and a rigid body. To achieve gait control of the assembled quadruped robot, a master computer is introduced and a four-level control strategy is proposed. A. Inchworm Robot Figs. 2a and 2b show the sketch and CAD model of the inchworm robot. The body of the inchworm is composed of two rigid feet and a compliant reed. The rigid feet are made of resin printed with a multi-material 3D printer (Object 260 Connex 3, Stratasys). The reed is made from manganese steel with excellent resilience. Each foot is attached with a servo and the servo is connected to the other foot with a cable. The 11682 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 17,2021 at 20:54:10 UTC from IEEE Xplore. Restrictions apply. servo motors are driven by signals with different frequencies from a custom-designed micro-controller and powered by a 3.7 V lithium-polymer battery. The driving mechanism of the inchworm robot is formed by the two servos and the compliance reed. The buckling motion of the inchworm robot is driven by the rotation of the servo motors and pulling of the cables. The restoring force is supplied by the compliant reed. Hooks are aligned at the bottom of the front and back feet to create asymmetric friction. The asymmetric friction force transfers the buckling motion into the crawl locomotion."
        ]
    },
    {
        "image_filename": "designv11_63_0002194_s12206-020-1201-5-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002194_s12206-020-1201-5-Figure7-1.png",
        "caption": "Fig. 7. Mounted location of sensors.",
        "texts": [
            " The planetary gearbox test rig is shown in Fig. 5. The test rig consists of drive motor, magnetic powder brake, speed and torque sensor, test gearbox and other main parts. The test gearbox is a single-stage NGW-11 planetary gearbox with a transmission ratio of 12.5, the specific structural parameters of test gearbox are shown in Fig. 6. During the experiment, four vibration acceleration sensors are arranged in the planetary gearbox case, and the specific mounted location of sensors are shown in Fig. 7. Sensors 1# and 2# are installed at the input shaft end, sensor 1# is horizontal, sensor 2# is axial direction, sensor 3# is installed at the top of the gearbox, sensor 4# is installed at the output shaft end and is axial direction, and all sensors are installed near the ring gear. In the life-cycle degradation experiment process, the input shaft speed of planetary gearbox is about 1000 rpm, and the load current of the magnetic powder brake is 1 A (about 340 Nm). The sampling frequency of the vibration signal is 20 kHz, each time signal acquisition lasts for 12 s, and it is collected once every 5 minutes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002730_978-981-16-0084-5_55-Figure51.2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002730_978-981-16-0084-5_55-Figure51.2-1.png",
        "caption": "Fig. 51.2 a Complete end-effector assembly, b detachable tip",
        "texts": [
            " A grasper is the first instrument that the surgeon will insert inside the abdomen during laparoscopic surgery. Figure 51.1 is the CAD model of atraumatic grasper with a double-action jaw. The end-effector consists of 12 components, including the shaft. The wristedmotion is actuated with the help of five stainless steel wires (F0.45 (mm)). The yaw part and the pitch part are responsible for producing the two wristed DoF. The range of motion of the wrist is \u00b145\u00b0 in both the perpendicular planes, and the grasper can open up to 90\u00b0. The ground part (Fig. 51.2) is completely detachable. The advantages of a detachable end-effector mechanism are: \u2022 Interchangeability: Specialized tips like grasper, scissors, etc. can be interchanged with various handles. \u2022 Cost lowering: Basic components like tip inserts and handles can be ordered separately. \u2022 Thorough cleaning: Partial detachability allows better cleaning, which was not possible earlier. \u2022 User-friendly assembly and disassembly as during surgeries only one component needs to be interchanged. \u2022 There is no need to unlock the instrument during ongoing surgery"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001865_012013-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001865_012013-Figure4-1.png",
        "caption": "Figure 4. Device for hardening thread",
        "texts": [
            " The essence of magnetic abrasive polishing is that the processed surface of the part or powder with magnetic and abrasive properties placed in a magnetic field, report a forced movement relative to each other. Metal removal is carried out as a result of the force action of the powder on the surface of the part and the specified relative movements (figure 3) [17]. The method of static-pulse processing of SPD consists in periodic pulse action on the loaded surface by a striker through a statically loaded waveguide, which allows more precisely controlling the distribution of microhardness and residual stresses at a significant depth of the hardened surface layer (figure 4) [15]. Modern power engineering (MPMB 2020) IOP Conf. Series: Materials Science and Engineering 963 (2020) 012013 IOP Publishing doi:10.1088/1757-899X/963/1/012013 These methods are based on the use of special equipment and tools for processing complex-profile surfaces, which limits their use in conditions of multi-product production. In view of this, it was proposed as a method of finishing processing to apply rolling toroidal roller of complex profile surfaces directly on metal-cutting equipment with CNC without changing the scheme of basing and fixing the workpiece"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002347_j.msea.2021.140880-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002347_j.msea.2021.140880-Figure13-1.png",
        "caption": "Fig. 13. Meshed structure with different surface roughness for (a) a smooth, (b) a medium-roughness, (c) and a high-roughness surface.",
        "texts": [
            " In Byun\u2019s model, R1 is the radius of the fillet (mm), and R2 is the radius of the corner (mm). R1 and R2 were set to 0.045 mm and 0.01 mm, respectively. In addition, Ra(\u03b8) was defined as 0 when \u03b8 is 0\u25e6 or 90\u25e6. Generally, the model could not predict the Ra(90) value because tan \u03b8 became infinite. Among them, Pandey\u2019s model showed a similar trend to the J. Lee et al. Materials Science & Engineering A 807 (2021) 140880 experimental results; however, a modification factor is required to employ the SLM method. Fig. 13 shows the meshed structure of three different surface roughness configurations. A rectangular-type mesh shape was used for the structural and thermal computational analyses. The total number of elements was 3024, 3210, and 3180 for the low, medium-roughness, and high-roughness specimens, respectively. Details of Ti\u20136Al\u20134V were provided for each specimen analysis. A fixed boundary condition was given on the narrow side surface for each structure, and a tensile force of 100 kN was applied on the other side for the mechanical analysis which used linear elastic model based on Hooke\u2019s law"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002935_iemdc47953.2021.9449597-Figure7-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002935_iemdc47953.2021.9449597-Figure7-1.png",
        "caption": "Fig. 7. Friction loss and approximated profile of the dynamic viscosity and density of the coolant based on the inlet (60\u2103) and outlet (112.5\u2103) temperature and pressure. Temperature is distributed linearly; and rotational speed is 30000 r/min. (a) bearingless motor, (b) SPM motor with AMBs alt. 2.",
        "texts": [
            " The combined negative position stiffness of AMBs and UMP of the SPM rotor give similar overall UMP forces as for the bearingless rotor. The AMB losses are computed for the 7 A bias current present in all coils, nominal rotor speed, and rotating control current vector with 20% magnitude of the maximum control current at 1st harmonic of 17000 and 30000 r/min. Similarly, for the bearingless motor rotor the suspension currents are at 20% of its maximum values. Friction loss distribution have been computed for the working fluid R1233zd(E) used as a coolant. The major losses are because of friction (Fig. 7 and 8). The nominal torque and motoring current have been used at electrical loss computation. Additional sinusoidal suspension current vector with rotational frequency is considered. In Fig. 8, losses for the cases vary with mechanical speed \u2126 and suspension current amplitude is,a conditions:  case 1: \u2126 = 30000 r/min, is,a = 0.2imax  case 2: \u2126 = 30000 r/min, is,a = 0  case 1: \u2126 = 17000 r/min, is,a = 0.2imax  case 2: \u2126 = 17000 r/min, is,a = 0 Table V lists key operational parameters for the studied cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002783_j.engfailanal.2021.105463-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002783_j.engfailanal.2021.105463-Figure6-1.png",
        "caption": "Fig. 6. DPF tube-gasket assembly (left) and assembly technique using push tool (right).",
        "texts": [
            " In this work it is investigated how well the leakage and the assembly force of a gasket are predicted with simulations. In this case the installation of a DPF tube in an exhaust aftertreatment system is studied in detail. This installation, and the measurement of the assembly force, is presented below. This gives a reference case for finite element simulations. During the DPF-tube assembly, the gasket is placed on the DPF-tube against a support ring that is welded to the DPF tube as presented in Fig. 6. The DPF tube with the gasket is pushed into the silencer wall using a pushing tool. The pushing tool is a disc welded to a number of stiffener supports, that are welded to a central cylindrical boss, as presented in Fig. 6. The tool is explained in detail in [20]. After pushing the DPF tube such that the clips attached to the DPF tube are completely inside the lid of the silencer wall shown in Fig. 7, the intermediate and the top cover are assembled (Fig. 3). With the position of the nut and the washer at the boss in Fig. 6, torque is applied to the nut to push the DPF inside. To check the force required to assemble the DPF tube assembly into the silencer wall, load measurements are conducted using a ring shaped load cell transducer placed between several washers and a nut as shown in Fig. 8. As one manually applies torque to the nut, the force required to push the DPF tube assembly a certain distance inside (which is around 1.6 cm, ensuring the 3 welded DPF clips are within the mark of the silencer lid), is sensed by the ring transducer and recorded"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002783_j.engfailanal.2021.105463-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002783_j.engfailanal.2021.105463-Figure2-1.png",
        "caption": "Fig. 2. Exhaust aftertreatment system geometry, with DPF filter disassembled, the lid (left), the DPF filter (mid) and the remaining exhaust aftertreatment system (right). The DPF filter has attachment brackets on one end that are bolted in the silencer, and is supported by the gasket (black) on the other end. The DPF-tube assembly bolts are of special type to work in high temperature.",
        "texts": [],
        "surrounding_texts": [
            "Engineering Failure Analysis 126 (2021) 105463\nfilter is presented in Figs. 2 and 3. Soot particles are trapped in the DPF filter, and thus the filter needs to be reconditioned regularly to remove excess particles. The installation in this case is thus equiped with a gasket to avoid leakage and to handle manufacturing tolerance deviations. The considered gasket geometry is presented in detail in Fig. 4.\nThe considered gasket connection requires simple assembly to reduce the time in production and workshop. For this reason, and to fit within the design space, the gasket is designed as a radial seal connection. In a radial seal connection the compression is applied on the outer surface of the gasket. The alternative is an axial face seal where the compression is applied at the cross section surface [6]. The assembly process of the considered DPF filter is only performed with virgin gasket specimens. The reason is that every time the DPF tube is cleaned, the gasket is replaced with a new one. The life of the gasket in service is approximately 6\u201312 months depending on the frequency of the DPF filter being cleaned.\nDuring operation of the vehicle, the temperature that the gasket is exposed to is maximum 500 \u25e6C. A graphite gasket is thus chosen because of its high deformability and low creep in the considered temperature range [7]. Above 500 \u25e6C, graphite gaskets could be subject to oxidation [8]. The graphite gaskets gained popularity when handling and trade with asbestos was recognized as work health hazard, with a sharp decrease of the use of asbestos in Europe during the 1990s [9]. Asbestos has otherwise excellent properties that could easily withstand up to 800 \u25e6C [10]. At higher temperatures it is also possible to use mica or vermiculite, that are basically clay minerals which can be compressed using roll forming with or without a binder. These gaskets are also structurally stable up to 800 \u25e6C [11]. However, graphite is preferred for this application, because of good compressibility and less complex composition. The considered DPF graphite gasket is reinforced with stainless steel wires that run in the form of a wire-mesh along the circumferential direction and at a 45 degree angle to the axial direction of the gasket. This wire is to be considered as a support for the graphite base material. The carbon content of the considered graphite base material is above 98 %. The gasket is presented in Fig. 5.\nStandards for testing of gaskets in flanged joints are reviewed in [12], where also a test matrix is proposed for acquisition of data for standard gasket designs and materials. The work includes different design parameters to characterize the gasket, such as maximum stress until damage, minimum assembly stress required and the minimum surface stress to keep the gasket contact leak tight. At elevated temperatures the loss of contact stress is also studied. The above mentioned attributes are considered to set design criteria for applications involving gasket flange connections. It is also observed that pure graphite at 400 \u25e6C can withstand a maximum stress of 100 MPa without getting damaged or experiencing extensive stress relaxation. To conclude, the work in [12] is an adequate baseline standard for comparison of the component testing in this work.\nLeakage of gaskets in elevated temperature is also studied in e.g. [13]. The conclusion is that the leakage in some cases decrease at elevated temperatures, due to increased deformation of the gasket material. Furthermore, the leakage rate depends very much on the surface roughness and surface non-tolerance [14]. This gives a challenge for truck exhaust aftertreatment systems, where the tolerances in the sealed connections are challenged by metal press forming and welding operations.\nIn other publications, further simulations are performed and compared with experimental observations of properties such as bolt pre-loading [15], nominal sealing effectiveness [16] and the effect of surface roughness [17]. It is generally concluded that it is necessary to model the non-linear behavior of the gasket base material for accurate predictions. The long term stress relaxation is considered for example in [18]. The material is modeled as a visco-elastic\u2013plastic material due to the strain-rate dependence of the material, so that the contact pressure could be investigated over time. A numerical simulation is also performed to optimize the bolting sequencing and pre-load on each bolt to predict the effect of creep relaxation and the loss of contact pressure.\nThe effect of pre-load and creep on the leakage through metal reinforced graphite sheet gaskets is predicted analytically using equations obtained empirically through test data in [19]. The conclusion is that at very high stress levels, large deformation and rotation of the surrounding components in the assembly contribute to leakage. In detail, for the applied pre-load, the effect of creep does not significantly affect the tightness of the bolt flange connection at 500 \u25e6C for 10000 h.\nThe main results presented in the following sections are the outcome of an investigation of the safety margin of the gasket installation in the exhaust aftertreatment system of a commercial truck, and thus to investigate of the behavior of a gasket influenced by manufacturing tolerance deviations. Another purpose is to study how well a finite element model predicts the assembly process and\nA. Ravikumar and A. Rietz",
            "Engineering Failure Analysis 126 (2021) 105463\nleakage. This is to enable simulation based engineering in future product development. The results are partly from a thesis project [20].\nMost finite element software for simulation of mechanical structures are based on a displacement formulation of the considered\nA. Ravikumar and A. Rietz",
            "Engineering Failure Analysis 126 (2021) 105463\nproblem. This means that displacements are the primary variables, and then the strains are calculated by definition from displacements and the stresses are calculated based on a material model. The material model required thus defines the stresses in terms of for example the strains and the displacements.\nIn this work we focus on the assembly of virgin graphite gaskets, and no deloading is considered. This means that elastic material models are investigated as a first attempt. The reason for choosing the elastic models is to limit the number of material model parameters if possible. Hyperelastic models are generally preferred for large strain problems, and often used to describe for example the material behavior of elastic isotropic materials like elastomers and rubber that undergo severe deformations or large strains. The stresses are assumed to be independent of strain-rate. With these models consequently, no unloading or relief can be predicted as mentioned. Some of the existing hyper-elastic models are polynomial models, reduced polynomial models, Arruda-Boyce models, Ogden models and Marlow models, and the specific model needs to be chosen from experience or with trial and error [21]. Several models are investigated in [20], and reduced polynomial models proved to be adequate for the virgin graphite gaskets. For some applications, specific finite elements for gasket components are developed that enable a specified compression behavior [22]. However, with these gasket specific finite elements, the performance in simulations with mechanical contact is limited, which makes the technique difficult to apply in the assembly process studied in this work [20].\nAll the hyperelastic models are defined through their constitutive strain energy formulation. The strain considered here is the right Cauchy-Green strain,\ndefined from the deformation gradient tensor F with components\nFij = \u03b4ij + \u2202ui\n\u2202xj (2)\nexpressed in the displacement field u and where \u03b4ij are the components of the identity matrix, \u03b4ij = 1 if i = j and \u03b4ij = 0 otherwise. The xj variables define local coordinates. For isotropic materials, the strain energy depends only of the invariants of the right Cauchy-Green deformation tensor (C). Now, let I1 be the first invariant of the right Cauchy-Green deformation tensor (C). In a hyperelastic reduced polynomial model of order k, the strain energy is defined by\nU = \u2211k\ni=1 Ci0(I1 \u2212 3)i (3)\nwhere k\u2a7e1 is an integer and Ci0 are parameters. From a given strain energy U of a hyperelastic material, the so called second PiolaKirchoff stress S is found\nS = \u2202U \u2202E\n(4)\nwhere\nE = 1 2 (C \u2212 I) (5)\nis the Green\u2013Lagrange strain tensor and I is the identity tensor. Expressed in components,\nSij = \u2202U \u2202Eij\n(6)\nA. Ravikumar and A. Rietz"
        ]
    },
    {
        "image_filename": "designv11_63_0000384_1.a34550-Figure10-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000384_1.a34550-Figure10-1.png",
        "caption": "Fig. 10 Change of optimal solution territory according to the trajectory of y- and z-axis thrust command.",
        "texts": [
            " When the trajectory of Fz and Fy escapes from the dashed-line rectangle, the solution determined by the boundary condition has no connection with k. Therefore, the constant value k can be chosen as another constant value during this period, and this change of k results in the change of the territory where the optimal solution exists. Finally, if constant value k is chosen appropriately outside the dashed-line rectangle, it is possible that Fz and Fy exist again in a new dashed-line rectangle where the optimal solution exists. For example, the trajectory of the thrustsFz andFy is marked with #1, #2, and #3 arrows in Fig. 10. The #1 trajectory starts at the origin Fy; Fz 0; 0 and exists within the dashed-line rectangle, and there is an optimal solution with the initial condition k FDsum\u22154 . Going to the #2 trajectory, there is a solution determined by the boundary condition. During #2 trajectory, the value of the constant k cannot be set again because FD3 is saturated with 0 and cannot be decreased less than 0. For the #3 trajectory, the value of the constant k can be set again so that there is an optimal solution while satisfying Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002753_s00202-021-01301-w-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002753_s00202-021-01301-w-Figure9-1.png",
        "caption": "Fig. 9 Magnetic flux lines distribution of in-wheel PM BLDC motor",
        "texts": [
            " 6 Effect of axial magnet segments to power of in wheel PM BLDC motor 1 3 has the highest efficiency. After that two-dimensional Finite Element Analysis carried out for the optimum designed 24 slot / 20 pole motor geometry and magnetic flux density is shown in Fig.\u00a08. The usage of magnetic core is determined by the proper flux distribution within specific limit values. The near saturated sections whose flux density is around 2\u00a0T are shown in the edges of teeth only. The magnetic flux lines distribution is homogeneous and smooth as given in Fig.\u00a09. The certain knowledge of the magnetic flux lines distribution taking these factors into account is essential for the accurate prediction of the motor performance. When the design parameters of the motor are determined properly, this specified magnetic flux line distribution should be homogeneous. One of the observations is that, increase in the number of circumferential magnet segments in axial direction resulted in lower eddy current losses and higher efficiency of in-wheel PM BLDC motor. Even if the rate of increase is low, this is important for motor designers and producers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002354_s42417-021-00283-0-Figure8-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002354_s42417-021-00283-0-Figure8-1.png",
        "caption": "Fig. 8 Radial static displacement of the output bearing seat and the inner bracket 1",
        "texts": [
            " When the output bearing seat is subjected to vibration, the vibration is transmitted to the inner bracket 1 through the contact surface. Figure\u00a07 presents the boundary conditions of the output bearing seat and the inner bracket 1. The contact type between the output bearing seat and the inner bracket 1 surfaces was set to bonded contact. The outer surface of the inner bracket 1 was defined as a fixed support, and a load F ranging from 100 to 500\u00a0N was applied to the inner surface of the output bearing seat along the Y direction. Figure\u00a08 shows the radial static displacement of the output bearing seat and the inner bracket 1 under load F, based on which, the deformation of the contact surface can be obtained. (10)zij = kij \u2212 2mij + i cij 1 2 (i, j = 1, 2, 3, 4, 5, 6) (11){X} = [z( )]\u22121{F}. (12) FT = kn(xr \u2212 xn) + cn(x\u0307r \u2212 x\u0307n) = [(kn + i\ud835\udf14cn)(Xre \u2212i\ud835\udf19r \u2212 Xne \u2212i\ud835\udf19n)]ei\ud835\udf14t (13)Tn = |||| FT F |||| = ||||| (kn+i cn)(Xre \u2212i r \u2212 Xne \u2212i n ) F0 ||||| 1 3 The deformation \u03b4 of the contact surface between the output bearing seat and the inner bracket 1 for different values of F is shown in Table\u00a02"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002630_iccia52082.2021.9403540-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002630_iccia52082.2021.9403540-Figure5-1.png",
        "caption": "Figure 5. error in x and y direction",
        "texts": [
            " The trajectory is defined by: ) 16 sin(3), 16 cos(3 tytx rfrf \u03c0\u03c0 \u2212== The initial error is [-10cm, -20cm, -0.2rad]. Authorized licensed use limited to: Carleton University. Downloaded on May 31,2021 at 12:05:31 UTC from IEEE Xplore. Restrictions apply. As shown from Figure 2 robot with robust LMPC, followed the reference trajectory with satisfactory performance. The control that comes from robust LMPC is shown in Figure 3 with handling the constraints. The robustness of the proposed controller has demonstrated in Figure 4 and Figure 5 which shows robust LMPC remains in the tube along the trajectory against LMPC results. It is shown in the LMPC case, multiple times the trajectory of error came out from the tube, especially after rejecting the initial error and when the disturbance free case, reached the steady state error. Data of the execution time came into Table 1; which is shown that according to the time step (Ts is 0.2s) and by imposing a constraint on max iteration in optimization problem [20], it does remain real-time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002454_ipemc-ecceasia48364.2020.9368059-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002454_ipemc-ecceasia48364.2020.9368059-Figure1-1.png",
        "caption": "Fig. 1. 9-slot/6-pole PM machines: (a) 9-slot stator; (b) conventional consequent-pole rotor; (c) Sin+3rd shaped rotor.",
        "texts": [
            " Otherwise, the multilayer winding and N-S- iron-S-N-iron sequence rotor cannot reduce the odd-order harmonics in CPM machine, which also lead to high torque ripple. The odd-order and even-order harmonics can both be suppressed by adopting the Sin+3rd shaped rotor in CPM machine. Therefore, this paper will investigate the influence of Sin+3rd shaped rotor on the torque ripple suppression of CPM machine. II. STRUCTURE OF MACHINES The 9-slot/6-pole surface-mounted consequent-pole permanent magnet machine is selected to investigate the influence of Sin+3rd shaped rotor on the torque ripple of CPM machine. As shown in Fig. 1(a), the stator is single layer concentrated winding of conventional 9-slot/6-pole PM machine. Fig. 1(b) shows the conventional 6-pole consequentpole rotor. Fig. 1(c) shows the Sin+3rd shaped rotor investigated in this paper. 1962 20 20 IE EE 9 th In te rn at io na l P ow er E le ct ro ni cs a nd M ot io n C on tro l C on fe re nc e (I PE M C 20 20 -E C C E A si a) | 97 8- 1- 72 81 -5 30 1- 8/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D O I: 10 .1 10 Authorized licensed use limited to: University of Exeter. Downloaded on June 03,2021 at 08:10:38 UTC from IEEE Xplore. Restrictions apply. Fig. 2 shows the shaping diagram of PM and iron rotor. The PMs are adopted Sin+3rd shaping and the salient rotor is shaped a certain width on both sides based on the minimum height of PMs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001251_0309324720936894-Figure9-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001251_0309324720936894-Figure9-1.png",
        "caption": "Figure 9. Moment 1: (a) proposed model and (b) FE model.",
        "texts": [],
        "surrounding_texts": [
            "Figure 8. Finite element model in ABAQUS. will lead to a decrease in computational speed, so this article selects the number of pieces as 20. Based on the above results, the mean error of contact force along axis is 7.1% solved by the sum of the error on each slice divided by the number of slices, and the mean error of contact force of one tooth is 4.2% solved by the sum of the error at each moment divided by the number of moment sampling points, as shown in Figures 13 and 14. The comparison results verified the proposed model and the result of the proposed model well reveals the pattern of contact force distribution of beveloid gear pair. Table 3. Status of each moment. Moment Rotation angle Mesh status Moment 1 28 Two teeth in contact Moment 2 68 One tooth in contact Moment 3 108 Two teeth in contact Figure 12. Effect of the number of pieces."
        ]
    },
    {
        "image_filename": "designv11_63_0000503_012020-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000503_012020-Figure2-1.png",
        "caption": "Figure 2. Pressure experienced by a mill on a water wheel",
        "texts": [
            " By using water-liquid fluid material, stating that the waterwheel moves based on water flow by selecting \"Relative to cell zone: fluid-solid\" means that the speed of this wheel will be adjusted to the conditions in the water flow. 3rd NICTE IOP Conf. Series: Materials Science and Engineering 725 (2020) 012020 IOP Publishing doi:10.1088/1757-899X/725/1/012020 Numerical analysis of the water wheel assumed that the waterwheel was half floating on the surface of the water, in accordance with the limitation of the problem that the incoming water flowing at a speed of 5 m/s then from the flow moves the windmill. The flow rate of the water that hit the blade on the water wheel causes the water wheel to rotate as shown in Figure 2 and Figure 3. In Figure 2 the red color is a waterwheel that gets pressure from the flow of water with a number of 12 pieces of blade. Where from the simulation results the pressure distribution experienced by each water wheel blade is evenly distributed, shown from the yellow color of each water wheel blade with an even pressure value of 43738.4 Pa. Figure 3 shows the results of the velocity distribution of water flowing on each blade of the waterwheel. The red color shows the highest flow velocity area on the waterwheel blade 3rd NICTE IOP Conf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001440_s00170-020-05863-0-Figure4-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001440_s00170-020-05863-0-Figure4-1.png",
        "caption": "Fig. 4 Contour plot of temperature distribution results at the middle of the substrate",
        "texts": [],
        "surrounding_texts": [
            "The increased speed of the reinforcement particles that displays a proper flow of molten liquid is brought about by the increase in laser power; also noting that an increased laser power causes improvement, standardized distribution, and spheroidization as shown in Figures 3, 4, and 5. As the laser beam moves from one point to another as shown in Figs. 3, 4, and 5, the temperature changes and the microstructures are determined by the laser input, scanning speed, and fast cooling rate. Figure 6 shows the microstructure of the titanium alloy base metal which acts as heat sink. High cooling rates form dendritic structures within the coatings. The major factors that determine the formation of the dendritic structure are the thermal gradients within the substrate during cooling and the cooling rates as shown in Fig. 7. The faster and slower cooling rates cause the spacing of the dendritic arm. Dendrite arm spacing has great influence on the mechanical properties of additive manufactured parts. The dendrite arm spacing is divided into primary and secondary. Crystal structures are modified by the rate of solidification in the microstructure. The distance from the Ti-6AI-4V alloy substrate determines the peak temperature distribution in the molten pool. There is a notable affiliation between the powder feed rate and laser scanning speed. The turbulent Marangoni convection current is the momentum and turbulence that occurs in the melt pool, which in fact is brought about by the blowing of the powders into the melt pool through the argon gas which carries the powders during laser metal deposition [20, 21, 26]. A decrease in the aspect ratio of columnar grains comes into effect by the increase in equiaxed grains as the build height increases. Processing conditions on grain structure with its influence were also noted. The microstructural development can also be affected by other processing parameters such as the powder flow rate as suggested by the results. Therefore, the microstructural development may have been conclusively determined by the input energy density, which could be defined by laser power, scanning speed, powder flow rate, and thermal history. Hence, there is influence on microstructure when any processing parameter affects the energy density and thermal history. The increase in line energies reduces its influence on the molten pool width, as confirmed by the simulations results. Only for various process parameter combinations were the molten pool depth determined by simulations, with constant line energy resulting in near constant melt pool depth. Higher line energies increased the molten pool depth. The temperature history within the material is prevailed by the process, which influences the mechanical properties of the fabricated coatings."
        ]
    },
    {
        "image_filename": "designv11_63_0000292_s40997-019-00330-y-Figure3-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000292_s40997-019-00330-y-Figure3-1.png",
        "caption": "Fig. 3 Dynamic analysis of active parts",
        "texts": [
            " The metamorphic process of mechanism based on augmented Assur group is the process of transforming 1-DOF augmented Assur group into basic Assur group. The 2-DOF metamorphic mechanism shown in Fig.\u00a01 consists of an augmented Assur group, and its four metamorphic configurations are shown in Fig.\u00a02. According to the above-mentioned structural theory and formation methodology of metamorphic mechanisms based on augmented Assur groups, the unified dynamics modeling method of constrained metamorphic mechanism is studied. The driving forms of active parts are shown in Fig.\u00a03. Figure\u00a03a and b shows active parts in pure rotational and pure prismatic form, respectively. According to the Newton\u2013Euler equation (referred to as N/E equation), the dynamic equations of active parts can be written as follows, where Fi\u22121,i and Fi+1,i are the force vectors at component Li exerted by components Li\u22121 and Li+1 , respectively, and (1) { Fi\u22121,i \u2212 Fi,i+1 + Fi = miC\u0308i Mi\u22121,i \u2212 li \u00d7 Fi\u22121,i \u2212Mi,i+1 \u2212 hi \u00d7 Fi,i+1 +Mi = IC,i\ud835\udf3ai Fi+1,i = \u2212Fi,i+1 ; Mi\u22121,i and Mi+1,i are the moments at component Li exerted by components Li\u22121 and Li+1 , respectively, and Mi+1,i = \u2212Mi,i+1 ; Fi is the external force vector acting on the component Li ; Mi is the external moment acting on the component Li ; IC,i is the moment of inertia of the component Li around the centroid Ci ; i is the angular acceleration of the component Li ; C\u0308i is the acceleration vector at the centroid of the component Li"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001304_icc40277.2020.9149416-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001304_icc40277.2020.9149416-Figure6-1.png",
        "caption": "Fig. 6. Low Altitude UAV Node Model",
        "texts": [],
        "surrounding_texts": [
            "In multi-UAVs network simulation scenario, we defined 4 kinds of node models, which are ground station, high altitude UAV, low altitude UAV, and terrestrial unit. 1) Ground Station Node Model Ground station node model is designed as a OSI protocol stack server equipped with one FDM wireless interface. Based on the high gain of the directional antenna, the ground server can connect with the high altitude UAVs via long-distance wireless FDM link. The ground station can handle traffic flows to and from remote UAVs and terrestrial units. 2) High Altitude UAV Node Model High altitude UAV node model is designed as a router with e interfaces. The IF-1 is an FDM interface, which is equipped with a PAA directional antenna. The IF-1 can provide high gain and long transmitting distance between high altitude UAV and ground station. The IF-2 is a TDMA interface, which can provide collision-free wireless links between the high altitude UAVs. The IF-3 is an 802.11 interface, which provide links between high altitude UAVs and low altitude UAVs. 3) Low Altitude UAV Node Model Low altitude UAV node model is designed as a two 802.11 interfaces router. The IF-1 is used to connect with high altitude UAVs and the IF-2 is used to connect with terrestrial units in the same area. With different BSS-ID configuration, the different 802.11 interfaces can work without interference. 4) Terrestrial Unit Node Model Terrestrial unit node model is designed as an 802.11 client with full OSI protocol stack. It can generate traffic flows in the application layer module and run routing protocols in the network layer module. With the 802.11 interfaces, terrestrial units in the same area can communicate each other directly or remotely via multi-hop forwarding. The low altitude UAVs that cover the area can provide relay ability for these terrestrial units. Authorized licensed use limited to: East Carolina University. Downloaded on July 29,2020 at 12:59:54 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_63_0000068_edpe.2019.8883869-Figure13-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000068_edpe.2019.8883869-Figure13-1.png",
        "caption": "FIGURE 13. Grid of stator.",
        "texts": [
            " And the efficiency reaches up to its maximum, namely 90%, when the speed is rated speed (300r/min,110Hz). From charts above, the motor designed in this paper has a good performance. The Block Lanczos method of Ansys software is adopted to analyse the structural modal of motor. The modal steps of solution and extension are both 40. Not considering the influence of motor frame, 3-D modal FEM analysis of stator core and winding is conducted. Grid, vibration pattern and vector diagram of radial natural modal of stator are shown in Fig.13 and Fig.14. Natural frequencies corresponding to each vibration patterns are listed in Table VI. 8542 VOLUME 4, 2016 Four test points are located on the motor, and acceleration sensor is placed at each test point. The NO.1 and NO.2 test points are located at the feet of motor, and NO.3 and NO.4 points are located at the top of motor. The inverter switching frequency is 2kHz. Acceleration of each point is tested at rated speed and load. Tested accelerations of four points under different frequencies are listed in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001118_10426914.2020.1772489-Figure5-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001118_10426914.2020.1772489-Figure5-1.png",
        "caption": "Figure 5. Gear tooth profile with a circular fillet.",
        "texts": [
            " A line is constructed from the involute profile limiting radius to a tangent point to the base circle. Two circles are drawn by considering line length as the radius for load side profile and host side profile of the adjacent tooth. A circle is constructed such a way that it is tangent to root circle and the two circles constructed as discussed above. The three circles and the tooth involutes profiles are shown in Fig. 4. The unwanted portions of the constructed lines are trimmed to get the complete gear tooth profile. The final profile is shown in Fig. 5. The gear parameters used in this study are given in Table 6. The AutoCAD DXF file is imported into the hyper mesh software. There the file is converted into neutral form (IGES). The IGES format of the profile has been imported to the machine to develop the WCEDM program. The gear blanks are case hardened and carburized with surface hardness value of 55HRC. Since the study of generated profile metrology is aimed at this work, only a few gear teeth were cut and inspected. For each case, two gears were prepared"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001916_icem49940.2020.9270720-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001916_icem49940.2020.9270720-Figure6-1.png",
        "caption": "Fig. 6: Construction of motor Linix 45ZWN24-40",
        "texts": [
            " Stator currents in abc reference frame iabc can be deformed by harmonic components in Back-EMF voltages eabc and by power inverter voltages uabc deformed by dead-time. Any misalignment between Back-EMFs eabc and stator currents iabc causes torque ripple in produced electromagnetic torque Te. Any misalignment between electromagnetic torque Te, load torque Tload and cogging torque Tcogg causes torque ripple in produced mechanical torque and thus also in speed. 1241 Authorized licensed use limited to: UNIVERSITY OF NEW MEXICO. Downloaded on May 15,2021 at 03:47:51 UTC from IEEE Xplore. Restrictions apply. Fig. 6 shows construction of 3-phase PMSM Linix 45ZWN24-40 and tab.1 contains its parameters. We tried to create the most accurate and the most reliable 2D model of this motor in Finite Element Method simulation software ANSYS-Maxwell. Then the FEM 2D model of the motor was coupled with simulation of Field Oriented Control in ANSYS-Simplorer(Twin Builder) as is shown in fig. 8 to simulate transient state. Fig. 7 shows 2D FEM model of used motor Linix 45ZWN24-40 in ANSYS-Maxwell software. Simulation model in fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0001765_ecce44975.2020.9236354-Figure2-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0001765_ecce44975.2020.9236354-Figure2-1.png",
        "caption": "Fig. 2. 2D schematic of investigated motor",
        "texts": [
            "00 \u00a92020 IEEE 1083 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 19,2021 at 14:42:00 UTC from IEEE Xplore. Restrictions apply. tion. A double layer winding is selected due to its reduced endturn length. The windings are wound around thermal plastic support structures in place of conventional laminated teeth as presented in Fig 1. The rotor has magnets in a Halbach configuration consisting of three segments per pole to circumvent the need for rotor core [1]. A non-magnetic rotor support is used to hold the magnet as shown in Fig 2. The coreless rotor minimizes the magnet loss and has zero rotor core loss. This helps to keep simple thermal management for the rotor. The considered machine is optimized to maximize the torque density and minimize the total loss while maintaining the parameters within the constraints given in Table I. Although the optimization is not the focus of this research, it was carried out to give the best platform for the proposed thermal concept evaluation. The simulated Back-EMF and torque profile are shown in Fig 3 and 4, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0002423_012128-Figure6-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0002423_012128-Figure6-1.png",
        "caption": "Figure 6. Scaled re-entrant cubic unit cell.",
        "texts": [
            " Tests are conducted with the combinations shown in the orthogonal array to identify the combination that provides the desired output. The optimal combination of the factors and levels were determined through analysis. Not all test cases could be analysed, as test cases 1,2,4 and 7 rendered an infeasible assembly for the combination of link length and link angle. A link length of 25mm and a link angle of 70\u00b0 provides the favorable output with a Poisson ratio of -0.576 and an induced stress of 228.99 MPa. Based on the findings the first prototype is fabricated (Figure 6). The proposed model has good scope to be scaled. This is possible by connecting just one link in the direction to be scaled as shown in Figure 6. ICRIET 2020 IOP Conf. Series: Materials Science and Engineering 1070 (2021) 012128 IOP Publishing doi:10.1088/1757-899X/1070/1/012128 Designing an auxetic structure that provides the maximum negative Poisson\u2019s ratio is always desired. But, fabricating such a complex geometry may at times be infeasible. Both, auxetic nature with manufacturing feasibility needs to be considered while designing the auxetic geometry. This article has proposed a new re-entrant cubic unit cell that has potential for mass production using conventional machines"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000377_012003-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000377_012003-Figure1-1.png",
        "caption": "Figure 1. 3D model of the parallel robot, with the seat and human operator models fixed to the mobile platform, where [9]: FP3 - the fixed platform; MP3 - the mobile platform; 1-6 - kinematical chains of SPS type; O - the origin of the fixed reference system; Cm - the center of mass of the human operator model; F - frontal point on the forehead of the human operator model; Oxyz - fixed reference system.",
        "texts": [
            " Both at fixed platform level and at mobile platform level, the centers of the spherical joints belonging to different kinematical chains are coincident 2 by 2, thus forming double spherical joints. The minimum distance between the centers of the 2 spherical joints of the same kinematical chain SPS is 800 [mm], the maximum distance is 1300 [mm], consequently the maximum stroke of the driving prismatic joint is 500 [mm] [9]. The 3D model of the parallel robot, obtained by using the SolidWorks tools, is presented in Figure 1. For the application of flight simulator, the 3D models of a seat and a human operator were considered [9]. The components have materials chosen from the SolidWorks database. The 3D model of the human operator has the height of 1.75 [m] and the mass of 75 [kg]. In order to obtain the proposed weight of this model, a custom material with the mass density of 1007.27 [kg/m3] was defined [9]. The simulation of parallel robot operation was accomplished in a SolidWorks Motion study. The gravitational acceleration of 9806",
            " International Conference on Applied Sciences Journal of Physics: Conference Series 1426 (2020) 012003 IOP Publishing doi:10.1088/1742-6596/1426/1/012003 The lengths L1\u00f7L6 between the centers of the spherical joints of the same kinematical chain, from 1 to 6, are presented in Table 1. International Conference on Applied Sciences Journal of Physics: Conference Series 1426 (2020) 012003 IOP Publishing doi:10.1088/1742-6596/1426/1/012003 The considered motions are combinations of translations along x and z axes and a pitch type rotation around y axis (according to the Oxyz reference system presented in Figure 1, with a corresponding tilt of mobile platform, as shown in Figure 3. In motion 1, the mobile platform back side is ascending, while in motion 2, the front side is descending. The total duration of motion is set to 0.5 [s]. As results, velocity and acceleration variations of the frontal point (F) and of the center of mass (Cm) of human operator are obtained, as shown in Figure 4. International Conference on Applied Sciences Journal of Physics: Conference Series 1426 (2020) 012003 IOP Publishing doi:10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_63_0000214_j.ifacol.2019.11.038-Figure1-1.png",
        "original_path": "designv11-63/openalex_figure/designv11_63_0000214_j.ifacol.2019.11.038-Figure1-1.png",
        "caption": "Fig. 1. The aerial manipulator and coordinate frames.",
        "texts": [
            " (2013) proved that an extension of the measurement vector by acceleration measurements results in a significantly improved performance for a manipulator, furthermore, the observers based on extended or unscented Kalman filter (UKF) show nearly the same performance. Due to the restriction that the vast majority of related researches presumed only UAVs without a manipulator, the force/torque estimation for an entire aerial manipulator system consisting of both a floating base and a manipulator should be researched. In this work, different observers from the above three groups are redesigned and evaluated for an aerial manipulator with a 3 DOF manipulator attached to a hexarotor with variable tilted rotors as shown in Fig. 1. The rest of this paper is organized as follows. In Section 2, the system model of the aerial manipulator is derived. Three different observer designs are submitted in Section 3. In Section 4, the different approaches are parameterized and evaluated in four test cases. Finally, Section 5 concludes the paper and gives remarks on future work. Nonlinear rench Observer Design for an Aerial anipulator Marek Wilmsen \u2217 Chao Yao \u2217 Micha Schuster \u2217\u2217 Shixiong Li \u2217 Klaus Janschek \u2217 \u2217 Institute of Automation, Technische Universita\u0308t Dresden, Germany (e-mail: {marek",
            " (2013) proved that an extension of the measurement vector by acceleration measurements results in a significantly improved performance for a manipulator, furthermore, the observers based on extended or unscented Kalman filter (UKF) show nearly the same performance. Due to the restriction that the vast majority of related researches presumed only UAVs without a manipulator, the force/torque estimation for an entire aerial manipulator system consisting of both a floating base and a manipulator should be researched. In this work, different observers from the above three groups are redesigned and evaluated for an aerial manipulator with a 3 DOF manipulator attached to a hexarotor with variable tilted rotors as shown in Fig. 1. The rest of this paper is organized as follows. In Section 2, the system model of the aerial manipulator is derived. Three different observer designs are submitted in Section 3. In Section 4, the different approaches are parameterized and evaluated in four test cases. Finally, Section 5 concludes the paper and gives remarks on future work. Nonlinear rench Observer Design for an Aerial anipulator Marek Wilmsen \u2217 Chao Yao \u2217 Micha Schuster \u2217\u2217 Shixiong Li \u2217 Klaus Jansc ek \u2217 \u2217 Institute of Automation, Technische Universita\u0308t Dresden, Germany (e-mail: {marek",
            " (2013) proved that an extension of the measurement vector by accelera ion measureme ts results in a significantly improved perform nce for a manipulator, furthermore, the observers based on extended or nscented Kal an filter (UKF) show nearly the same performa ce. Due to the restriction that the vast majority of related res arches p umed only UAVs wi hout a manipulator, the forc /to que estimation for an entire aerial m nipulator sys em consisting f b th a floating b se and a m nipulator should be researched. In this work, different observers from the a ove three groups are redesigned and evaluated for an aerial manipulator with a 3 DOF manipul tor attached to a hex rotor with variable tilted rotors as shown in Fig. 1. The rest of this paper is organized as follows. In Section 2, the ystem model of the aerial m nipulator is derived. Thre different observer d signs are submit ed in Section 3. In Section 4, th differ nt app oaches ar parameterized and evaluated in four test cases. Finally, Section 5 concludes the paper and gives remarks on future work. li W s si f i l M i l t arek il sen \u2217 hao ao \u2217 icha Schuster \u2217\u2217 Shixiong i \u2217 laus Janschek \u2217 \u2217 Institute of uto ation, echnische niversita\u0308t resden, er any (e- ail: { arek",
            " (2013) proved that an extensio of the measurement vector by acceleration measurements results in a significantly improved performance for a manip lator, further ore, the observers based on extended or unsce ted Kalman filter (UKF) show nearly the same performance. Due to the restriction that the vast majority of related researches presumed only UAVs without a manipulator, the force/torque estimation for an entire aerial manipulator system consisting of both a floating base and a manipulator should e researched. In this work, different observers from the above three groups are redesigned and evaluated for an aerial manipulator with a 3 DOF manipulator attached to a hexarotor with variable tilted rotors as shown in Fig. 1. The rest of this paper is organized as follows. In Section 2, the system model of the aerial manipulator is derived. Three different observer designs are submitted in Section 3. In Section 4, the different approaches are parameterized and evaluated in four test cases. Finally, Section 5 concludes the paper and gives remarks on future work. As shown in Fig. 1, let {I} be the inertial reference frame. {B} is the body fixed frame placed at the multicopter\u2019s center of mass. The link frames {Li} are placed in the joint of the manipulator with the assumption that the joint motors contribute the major part to the links mass and inertia. The position of the UAV\u2019s center of mass in {I} is represented by Ip = [px py pz] , which is the origin position of {B} described in {I}. Euler angles I B\u03c6 = [\u03d5 \u03d1 \u03c8] with a yaw-pitch-roll sequence are chosen as representation of the UAV\u2019s attitude"
        ],
        "surrounding_texts": []
    }
]