[
    {
        "image_filename": "designv11_30_0002924_s1063771020020104-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002924_s1063771020020104-Figure1-1.png",
        "caption": "Fig. 1. The arrangement of arrays of ultrasonic radiators. Levitating particles ordered in the rectangular grid are depicted in the center.",
        "texts": [
            " It is shown that levitation in the system is possible not only for small particles but also for linear segments whose lengths greatly exceed the wavelength, as well as for plane structures assembled of rectilinear segments joined either at a right angle or at an angle of 45\u00b0. It is proposed to stabilize the positions of particles levitating in air due to the formation of a grid of points distributed in the plane as local minimums of the Gor\u2019kov potential. The corresponding distribution of the acoustic field is provided by interference of fields of oppositely directed arrays of radiators as the field is focused near the levitation plane. The arrays of the ultrasonic radiators are arranged according to the scheme depicted in Fig. 1 and denoted by the letters A, B, C, and D. Every array contains 320 elements placed in a hexagonal grid of 16 rows by 20 elements in each row. The opposite arrays are focused into a common line that passes parallel to the arrays and equidistantly between them at a given height z0. The acoustic pressure field generated by the focused arrays is specified by the expression (1) ( ) ( ) ( )( ) ( ) ( )( )  \u2212 \u2212 \u2212 \u2212  \u2212    \u2212 \u2212 \u2212 \u2212  \u2212   2 2 0 0 0 2 2 0 0 0 exp ' ' 'exp , ' n n n n n m m m m m p ik x X + z Z P = p ik y Y + z Z+ r R r r R r R r R this method of focusing of opposite arrays cylindrical where are coordinates of radiators in arrays oriented parallel to the yOz plane (arrays A and C); are coordinates of radiators in arrays oriented parallel to the xOz plane (arrays B and D); n is the radiator number from 1 to 640 in arrays A and C, m is the radiator number from 1 to 640 in arrays B and D; is the acoustic pressure created by an individual radiator at unit distance; is the coordinate of the focusing line of the yOz arrays along the x axis; is the coordinate of the focusing line of the xOz arrays along the y axis; and z0 is the focusing height for all the arrays",
            " The particles move downwards under the action of the force of gravity; at the edges they move downward to a greater extent due to deterioration of the field localization (Fig. 5). The results of numerical simulation in the COMSOL Multiphysics environment demonstrate the possibility to provide levitation of particles in a rectangu- lar grid in a plane based on fields of four monochromatic cylindrically focused radiators. The position of levitating particles according to the numerical simulation corresponds to the theoretically predicted position. It should be noted that if the radiating arrays are arranged as shown in Fig. 1, multiple reflections from the array planes will occur, which will distort the field of primary waves. To minimize the field scattered by the arrays, it is proposed to place the radiators at a depth of a quarter of the wavelength in the holding plate. Then, the field scattered from the sensor body will be added in antiphase with the field scattered on ACOUSTICAL PHYSICS Vol. 66 No. 2 2020 the surface of the holding plate, which minimizes the amplitude of the wave scattered by the array in the levitation region",
            " Thus, the numerical estimate demonstrates that the primary field exceeds the field scattered by arrays by 15 times on average. Therefore, one can neglect the field scattered by the opposite array of radiators. ( ) ( ) ( )( ) ( ) 2 1' , 4 , exp , 2 exp . x z x z P k k P x z i M x z ik x ik z dxdz \u221e \u221e \u2212\u221e \u2212\u221e = \u03c0 \u03c0\u00d7 \u00d7 \u2212 \u2212   ( )( )exp , 2 i M x z\u03c0 ( ), 0.M x z = EXPERIMENTAL INVESTIGATIONS To experimentally verify the possibility of levitation and control for an ordered group of particles, a facility of four arrays of ultrasonic radiators (320 elements in each array) was designed according to the scheme depicted in Fig. 1. The elements in the arrays are arranged in a hexagonal grid with a step d = 11 mm in 16 rows and 20 radiating elements in each row. Radiating elements in a row are parallel connected and joined to the output of the harmonic signal amplifier. The entire system of the arrays is controlled using an STM32F407 microcontroller that implements a 64- channel 1-bit digital-to-analog converter. Binary signals arrive at TDA7297 amplifiers (the supply voltage is 12 V) whose outputs are connected to rows of the radiating arrays (16 outputs for each of the four arrays)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure6-1.png",
        "caption": "Fig. 6 Physical model of driving system on servo press.",
        "texts": [],
        "surrounding_texts": [
            "The control strategy of servo driver application for the press mechanism is illustrated in Fig.4. \u03b1, , \u2217and are the current angle of crank, the calculation output of Lagrange Equation, the reference torque of PTC and current speed of IM, respectively. The Predictive Torque Control (PTC) algorithm is explained and evaluated in this part. According to the basic equation of an IM, it is quite easy to describe the system with complex numbers and obtain a more compact system representation in the form + = + ( \u2212 ) (13) + = + (13) where \u03c4\u03c3 = \u03c3Lsr\u03c3 and r\u03c3 = Rs + kr 2Rr with kr = Lm Lr , \u03c4r= Lr Rr and \u03c3 = 1 \u2212 Lm 2 LsLr . After discretization with sampling time , the stator current can be calculated as is(k+1) = 1- Ts \u03c4\u03c3 is(k) + Ts r\u03c3\u03c4\u03c3 \u03c5s(k) + KrTs r\u03c3\u03c4\u03c3 1 \u03c4r -j\u03c9e1 \u03c8r(k) (14) The mechanical torque can then be predicted Tm(k+1) = 3 2 p (\u03c8s(k+1) \u00d7 is(k+1)) (15) Finally, the stator flux magnitude can be calculated by \u03c8s(k+1) = \u03c8s\u03b1 2 (k+1)+\u03c8s\u03b2 2 (k+1) (16) Hence, an -norm(linear) cost function can be defined as gh = T*-Tm(k+1) +\u03bb\u03c8 \u03c8s *-\u03c8s(k+1) (17) The torque reference value is generated by a conventional PD controller of computed torque, so T*= \u03c4 nratio + JM\u03b1dnratio (18) where nratio is the ratio of Gear box, and JM is total inertia of Gear box and motor. IV. SIMULATION RESULTS To check the performance of PTC, simulation test is firstly conducted on a 30kW, 948rpm IM. The electrical machine parameters are presented in TABLE \u0399. TABLE I. MACHINE PARAMETERS Parameter Value nominal power 30Kw nominal flux 1.4Wb nominal speed 948rpm nominal torque 302 \u2219 \u03c9nom synchronous frequency 33.3Hz stator resistor 0.137(20\u2103) Rr rotor resistor 0.163(20\u2103) Ls stator inductor 1.38mH rotor inductor 1.14mH mutual inductor 47.8 mH Crank trajectory tracking and desired motion PD+Computed torque control PTC+IM Gear box: Crank, Link, and Slide T*Calculate ref-torque e 3065 Authorized licensed use limited to: Carleton University. Downloaded on May 27,2021 at 10:48:34 UTC from IEEE Xplore. Restrictions apply. Parameter Value p pole pairs 2 Ts sampling time 100\u03bcs Jm Inertia 0.246 kg \u2219 m Fig.7 shows the torque reference and actual torque, reference speed and real feedback speed. The steady-state characteristics of IM at the reference speed of \u00b1948rpm are shown. The dynamic performance of a machine is analyzed with slope changes in reference speed from \u2212948rpm to 948rpm. The step changes in load torque are performed from zero to -302N\u00b7m at 0.6s. And the load repeats the torque 302N\u00b7m at 1.5s. It is observed that motor torque is tracking load torque with low ripple. The simulation of PTC of IM with Computed torque for servo press is shown in Fig.8. The dynamic preformace of the machine is examined from 0 rpm to about 335rpm while the magnatic flux of stator is steady about 1.4wb. As shown in Fig.8, the angle and speed of crank are observed for the proposed scheme indicating a good tracking. The errors of crank angle and speed are even smaller than 0.4%. These simulation results show proposed control strategy is better than conventional method in which our research group have adapted synchronous machine before [2]. These results indicate effectiveness of this method of servo driver used computed torque and FCS-MPC. V. CONCLUSIONS This paper presented a computed torque method based MPC for trajectories of servo press. The predictive control uses the optimal response of torque FCS-MPC loop to reduce the work of tuning PID parameters by hand. The control has 3066 Authorized licensed use limited to: Carleton University. Downloaded on May 27,2021 at 10:48:34 UTC from IEEE Xplore. Restrictions apply. high dynamic performance and small steady error respect to traditional industrial control method. For the future studies, an adaptive control and torque ripple can be incorporated into this control design."
        ]
    },
    {
        "image_filename": "designv11_30_0003476_0954408920948683-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003476_0954408920948683-Figure2-1.png",
        "caption": "Figure 2. Top tendon guide of continuum structure.",
        "texts": [
            " The parameters defined as below: ri Projection of the driving wire to the YAOAZA plane L Length of continuum structure (L\u00bc 280mm) Li Length of the driving wire DLi Variation of the driving wire hi Distance from Li to ZA-axis qi The radius of curvature of Li h Bending angle of continuum structure u Rotation angle of continuum structure a Rotation angle of turntable hT Displacement of vertical sliding table qT Displacement of horizontal sliding table The continuum structure is driven by three wires, which are fixed to I, II and III position of the top tendon guide, and cross through the hole of other guides, as shown in Figure 2. In order to establish the relationship between the variation of the driving wires and the position of the end of continuum structure, the parametric kinematics analysis was performed by using constant curvature mode,27,28 as shown in Figure 3. Base coordinate system (A) x\u0302A; y\u0302A; z\u0302A is the original coordinate system. Rotation coordinate system (B) x\u0302B; y\u0302B; z\u0302B is obtained by rotating of the A coordinate system around the ZA-axis for u degrees. System (C) x\u0302C; y\u0302C; z\u0302C is the free end coordinate system, the angle between XC-axis and XA-axis is h",
            " Free end rotation coordinate system (D) x\u0302D; y\u0302D; z\u0302D is obtained by rotating of the C coordinate system around the ZC-axis for u degrees, and the XD axis is parallel to the XA-axis. The relationship between the coordinate of the end point of the continuum structure in three-dimensional space and its rotation angle and bending angle can be deduced by using the trigonometric function xD yD zD 2 4 3 5 \u00bc L h \u00f01 Cosh\u00de Cos/ \u00f01 Cosh\u00de Sin/ Sinh 2 4 3 5 (1) According to the fixed position of the driving wires at the top segment of continuum structure, as shown in Figure 2. The variation of driving wire is analyzed when the bending angle is h and the rotation angle is u. According to the right-handed Cartesian coordinates, the positive direction is defined. Using the arc length theorem and the trigonometric function, the following results can be obtained DL1 \u00bc h h1 Sin\u00f0a /\u00de DL2 \u00bc h h2 Sin\u00f0a\u00fe b /\u00de DL3 \u00bc h h3 Sin 3p 2 / 8>>< >>: (2) Thus, the drive of the continuum structure can be achieved by controlling the variation of the wire by the servo motor. In order to obtain the workspace of the picking robot, the transformation matrix is used to analyze the forward kinematics solution of the picking robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002499_iecon.2014.7049073-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002499_iecon.2014.7049073-Figure9-1.png",
        "caption": "Fig. 9. Average flux path in a 4-pole induction machine.",
        "texts": [
            " Initially calculate the specific losses by dividing to stator steel weight (W/kg), where the rotor weight is neglected since most of the iron losses occur in the stator. Afterwards calculate the volt-amps per kg from ( )( )22 1 1 1 1 1 3 3VI kg M iron iron I E R I I E Wt Wt \u2212 = = (11) where EM is the voltage across the total inductance LM =L1+Lm and Wtiron is the stator steel weight. The calculation the field-strength H from Ampere\u2019s law over one pole flux path is now required. In a 4-pole machine the total length of the path (Fig. 9) can be approximated to ( ) ( )2 2 Poles syoke Ryoke mmf syoke Ryoke R R L R R \u03c0 + = + \u2212 (12) where RSyoke and RRyoke are taken as radii in the middle of stator yoke and rotor yoke. The current and volt-amps are ( )2 Conductors / Poles mmf PathsH L P I \u00d7 \u00d7 = (13) [ ] 2 2 PolesVI 32 2 stk mmf f HB L kL\u03c0 \u03c4 \u03c0 = \u00d7 \u00d7 \u00d7 \u00d7 (14) where f is the frequency, \u03c1 is the steel density, Lstk is the axial length, \u03c4 is the polar arc, and k is the tooth width/slot pitch ratio. From (13) and (14) the flux-density can be estimated from 19 2 Poles M stk mmf I E B f HL kL\u03c4 = \u00d7 \u00d7 (15) By plotting the specific losses in W/kg against the fluxdensity from the test data at various frequencies, we can estimate the core loss coefficients for any suitable core loss variation law, e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure10-1.png",
        "caption": "Fig. 10. Detail view of the fabricated RFECT system",
        "texts": [
            " It is difficult to implement RFECT technology inspection tools for large-diameter pipelines because many receiving sensors are necessary to ensure the full circumferential coverage of the pipeline, and consequently, the same number of LIA (Lock-In Amplifiers)s with receiving sensors, which measure the phase shift of the measured magnetic wave signal in the remote field zone due to a change of wall thickness, is required to implement an RFECT system for large-diameter pipelines. Implementing the parallel type of multi-channel digital LIA, a 16 in of RFECT system was fabricated like Fig. 10. As shown in Fig. 10, the sensors and exciter modules were equipped with all of the electronic systems for signal acquisition, processing and storage. The processed signal from all sensor channels can be transported by a sub controller which communicates with the control station by wireless communication in the high pressure of gas pipeline. The integrated PIBOT is shown in Fig. 11. Functions of each module were checked by examination at UPSF (Unpiggable Pipeline Simulation Facility). The UPSF was constructed to undertake several kinds of performance test about defect detectability, driving efficiency and pressure resistance ability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003978_978-3-319-14705-5_12-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003978_978-3-319-14705-5_12-Figure11-1.png",
        "caption": "Fig. 11 Theoretic trajectory of circular motion of BHQ-1",
        "texts": [
            " (26) we can get l j = \u2212 f \u00b7 r mg (29) \u03b2 (t) = arcsin l j L = arcsin ( \u2212 f \u00b7 r mgL ) (30) Thus when spherical robot BHQ-1 moves along a straight trajectory with a constant velocity the driving moments of two motors are { M1 = mg l j = \u2212 f r M2 = 0 (31) The simulation results of BHQ-1 moving straight are shown in Fig. 10, where the dashed line is the planned trajectory in theory or the target trajectory and the solid line is the planned trajectory by adding 1 % noise disturbance. From the simulation we can conclude that BHQ-1 can realize motion along a linear trajectory with the deduced dynamic model. Assume spherical robot BHQ-1 moves along a circular trajectory from the initial configuration to the final configuration, as shown in Figs. 11 and 12. In Fig. 11, a frame {oi xi yi } is established on the geometric center of BHQ-1 with its axis xi pointing to the center of the circular trajectory and its axis yi pointing to the tangential direction of the circular trajectory. With similar derivation to that of linear trajectory planning, we can deduce f 0 1 = 0, f 0 2 = 0, m0 3. Let p = 7mr2 5 and substitute the above variables for those Eq. (23), we can get \u23a7\u23aa\u23a8 \u23aa\u23a9 m0 1 = p\u03c9\u03071 m0 2 = p\u03c9\u03072 \u03c9\u03073 = 0 (32) Because m0 1, m0 2 are projections of the principal moment on axes i and j, it\u2019s easy to get \u23a7\u23a8 \u23a9 m0 1 (t) = mg \u21c0 l y + \u21c0 jy \u00b7 r m0 2 (t) = mg \u21c0 lx + \u21c0 fx \u00b7 r (33) where, \u21c0 lx , \u21c0 l y are the projections of length L on axes x and y respectively, \u21c0 fx , \u21c0 fy are the projections of the friction force f on axes x and y respectively. From Eqs. (16), (32) and (33), we can get \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 \u21c0 lx = 1 mg ( p r x\u0308 \u2212 \u21c0 fx \u00b7 r ) \u21c0 lx = 1 mg ( \u2212 p r y\u0308 \u2212 \u21c0 fx \u00b7 r ) (34) Through coordinates transformation we can get \u23a1 \u23a3 lxi lyi 0 \u23a4 \u23a6 = R(z,\u2212\u03b1) \u00b7 \u23a1 \u23a3 lx ly 0 \u23a4 \u23a6 = \u23a1 \u23a3 cos(\u2212\u03b1) \u2212 sin(\u2212\u03b1) 0 sin(\u2212\u03b1) cos(\u2212\u03b1) 0 0 0 1 \u23a4 \u23a6 \u00b7 \u23a1 \u23a3 lx ly 0 \u23a4 \u23a6 (35) where, \u03b1 is the angle that spherical robot BHQ-1 has moved along the circular trajectory from origin o (as shown in Fig. 11), lx , ly are the norms of \u21c0 lx , \u21c0 l y respectively, \u21c0 lxi , \u21c0 l yi are the projections of length L on axes xi and yi respectively (shown in Fig. 11), lxi , lyi are the norms of \u21c0 lxi , \u21c0 l yi respectively. From Eqs. (34) and (35) we can obtain { lxi = p mgr (x\u0308 cos \u03b1 \u2212 y\u0308 sin \u03b1) lyi = \u2212 p mgr (x\u0308 sin \u03b1 + y\u0308 cos \u03b1) + f \u00b7r mg (36) Because \u03b1 = \u03c9t , the driving moment of motor 1 can be got as M1 = mg \u21c0 l j + m\u03b2\u0308y L2 = \u2212 p r (x\u0308 sin \u03b1 + y\u0308 cos \u03b1) + m\u03b2\u0308y L2 (37) where, \u21c0 l j is the projection vector of L on axis j, \u03b2y is the angle that the mass deviates from axis k measured on plane o\u2032k j (as shown in Fig. 12). The driving moment of motor 2 is M2 = mg \u21c0 li + m\u03b2\u0308x L2 = p r (x\u0308 cos \u03b1 \u2212 y\u0308 sin \u03b1) + m\u03b2\u0308x L2 (38) where, \u21c0 li is the projection vector of L on axis i, \u03b2x is the angle that the mass deviates from axis k measured on planes o\u2032ki (as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001878_tmag.2015.2489701-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001878_tmag.2015.2489701-Figure4-1.png",
        "caption": "Fig. 4. Coil wrapped by using OGWM (left) and the PEC model (right).",
        "texts": [
            " Since both formulas (1) and (3) derived under different physical insights have used the same approximation, the deviations introduced can be reduced by using the average value of both (K 1 + M1)/2. Thus, a modified formula for OGWM can be obtained. For the capacitance wrapped by using OCWM, the formula of Massarini has better accuracy than the formula of Koch [5]. Therefore, (2) will be used for the further investigation of the capacitance for OCWM. The accuracy of formulas for both winding methods is validated in Section III. To reduce the mesh effort and computing time, a PEC model, instead of a complete coil model, is presented in Fig. 4. The red surface of each turn has a constant electric potential denoted with the yellow number. Moreover, the electric potentials along the turns are assumed to be linear [6]. Then, each wire with a circular cross section can be replaced by a rectangular surface with the same electric potential. The distance d between two layers has remained unchanged. Furthermore, \u03b5equi between two flat surfaces is required to ensure the capacitance of this model equal to the value of the original one\u2019s. For OGWM, the \u03b5equi1 value of a coil with m layers can be calculated as follows: \u03b5equi1 = d \u03b502rn (ClK 1 + ClM1 ) 2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure24-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure24-1.png",
        "caption": "Figure 24. Speed field contours for 900 rpm.",
        "texts": [],
        "surrounding_texts": [
            "The first stage of the iterative algorithm for solving the problem of optimizing the kinematic and geometric characteristics of the field stripper and stationary installations has been developed. The task of studying the movement of the grain mass is divided into two stages. At the first stage, a model of air mass movement in the considered installations is built and based on a mathematical model that takes into account the turbulence of the movement using the finite volume method; the field of velocities and pressures is calculated. The influence of the technological parameters of the installations on these ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 fields was studied: the diameters and rotation speed of the stripping drum and the additional drum-fan, the shape of the surface of the upper deck of the installation chamber. The calculations have shown that additional field experiments are required to build an adequate model of the stationary installation. A series of calculations was carried out according to the developed algorithm; the geometric and kinematic parameters of the installations were obtained."
        ]
    },
    {
        "image_filename": "designv11_30_0001865_cdc.2015.7403320-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001865_cdc.2015.7403320-Figure1-1.png",
        "caption": "Fig. 1. Magnetic field of a dipole in a 2D surface",
        "texts": [
            " By bringing any magnetic sensor close to the engine cylinder, the position of the piston inside the cylinder can be estimated. No instrumentation inside the engine cylinder is needed. All other sensing mechanisms need either embedded instrumentation inside the cylinder or a line of sight access to a device attached to the piston of the cylinder. Thus, the magnetic sensing principle enables a fundamental new method of position estimation. In many applications, a small magnet can be modeled as a dipole as shown in Fig. 1. The red arrow represents the direction of the dipole. According to the physics of magnetism, the planar components of the magnetic field generated by the element dm0 at the point S are given by [5]: Br = \u00b50dm0 2\u03c0r3 cos\u03b1, B\u03b1 = \u00b50dm0 4\u03c0r3 sin\u03b1 (1) Based on the assumption that a large ferromagnetic object is uniformly magnetized, its magnetic field can be obtained through the integration of the dipole model in a 3D space. For example, consider the magnetic field along the X-axis of the cylinder shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003792_iros45743.2020.9341529-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003792_iros45743.2020.9341529-Figure2-1.png",
        "caption": "Fig. 2: Overview of our system.",
        "texts": [
            ", a surgical robot with a chopsticksshaped end effector may effortlessly switch between grasping, moving and rotating various shapes of tissues with different deformability without switching hardware [17]. Importantly, human familiarity with chopsticks opens up the possibility of easily collecting expert demonstrations. Despite the wide range of applications involving chopsticks, we are not aware of any prior efforts to learn from human demonstrations for chopstick-based robot manipulation tasks. To this end, we analyze how different interfaces and user expertise levels affect the quality of demonstrations by comparing three data-collection methods (Fig. 2a): normal chopsticks (\u201cChop\u201d), motion captured chopsticks (\u201cMoChop\u201d), and a teleoperation interface (\u201cTeleChop\u201d). We conduct a within-subjects user study with 25 participants and examine how human factors affect the success rate of picking up everyday-life objects. One key finding is that user preferences for input modalities may not necessarily correlate to the quality of the resulting demonstrations. Although users rated teleoperation as unnatural and uncomfortable, they exhibited comparable performance using TeleChop, MoChop, and Chop",
            " Humans successfully teleoperated our robot to complete a challenging manipulation task: picking up a slippery glass ball with slippery metal chopsticks, without haptic feedback. We believe that teleoperation could be the preferred interface to yield on-hardware demonstrations for robots to learn from [24]. We designed a pair of chopsticks for motion-capture (\u201cMoChop\u201d) and an interface for users to teleoperate a robot holding chopsticks (\u201cTeleChop\u201d). Both were adapted from normal chopsticks (\u201cChop\u201d). See Fig. 2a. All methods used the same consumer-grade titanium chopsticks that differ only in colors. We 3D printed five light-weight, ball-shaped markers, wrapped them in reflective material and mounted them onto MoChop. To ensure users could still hold the chopsticks, we placed the markers near the tips and tails of chopsticks at different positions on each stick. Note the markers changed the weight distribution and added constraints to finger positions, e.g. users cannot hold the chopsticks at the tail. To track the motion of MoChop held by users, we used the OptiTrack motion capture system [25] with 11 cameras (Fig. 2b). The system uses optical reflection to track the position of markers. Its tracking error is up to 0.4mm, and the tracking updates at 100Hz. From the tracked markers\u2019 positions, we extracted MoChop\u2019s pose. We custom built a 6-DOF robot manipulator, assembled from components provided by HEBI Robotics [26]. We attached the TeleChop to this robot\u2019s end-effector. TeleChop has an actuated pair of chopsticks. The two chopsticks operated on the same plane: one was fixed at the actuator\u2019s body, and the other is attached at the actuator\u2019s output shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002769_ls.1499-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002769_ls.1499-Figure4-1.png",
        "caption": "FIGURE 4 Exploded view comparing the previous and current assembly solutions",
        "texts": [
            " With the improved assembly, the \u201c10-Drive spindle\u201d is in full contact with the rotating race of the rolling bearing; furthermore, it is also much thicker. This spreads the axial load over the upper race more evenly, as it would in a real application. For this reason, the material of the rolling bearing driving element is a steel alloy for the thrust roller/ball bearings (51107, 81107 TN). For the other rolling bearings types that can be tested, Table 3, the rolling bearing driving element can be made of an aluminium alloy for improved heat transfer, Figure 3. Figure 4 shows an exploded 3D view of the previous and current assemblies. This exploded view also shows the assembly order of both assemblies. Despite having more complex parts, the proposed solution is as difficult as the previous one to assemble. At this point, it should be apparent that in Figures 2 and 4A, the assembly is shown with a thrust roller bearing (81107 TN), but in Figures 3 and 4B the new assembly is shown with a tapered roller bearing (32008 X/Q). In fact, with the proposed assembly rolling bearings with dimensions up to the dimensions of the 32008 X/Q tapered roller bearing can be tested"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002948_s40435-020-00624-z-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002948_s40435-020-00624-z-Figure4-1.png",
        "caption": "Fig. 4 Top view",
        "texts": [
            "Hence, the rotation matrix is introduced as follows: R = Rz(\u03c6)Ry(\u03b8)Rx (\u03c8) = \u23a1 \u23a3 c\u03c6c\u03b8 c\u03c6s\u03b8 s\u03c8 \u2212 s\u03c6c\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 s\u03c6c\u03b8 s\u03c6s\u03b8 s\u03c8 \u2212 c\u03c6c\u03c8 s\u03c6s\u03b8c\u03c8 + c\u03c6s\u03c8 \u2212s\u03b8 c\u03b8 s\u03c8 c\u03b8c\u03c8 \u23a4 \u23a6 (2) where c\u03c6 = cos(\u03c6), s\u03c6 = sin(\u03c6), c\u03b8 = cos(\u03b8), s\u03b8 = sin(\u03b8), c\u03c8 = cos(\u03c8), s\u03c8 = sin(\u03c8). yAi Let one end of the pulley be attached at the exit points Ai and a contact pointCi at another, as shown in Fig. 3. The pulley moves with respect to its own coordinate system, denoted as Fai of axes xai , yai , zai . The angle produced between these two points is denoted as \u03b2i . The position of contact point Ci is denoted by position vector bci = [ cx cy cz ]T . The pulley can be rotated about zai axis of an angle \u03b3i as shown in Fig. 4. Therefore, the vector of cable lengths bli can be determined as follows: bli =b ci \u2212b p \u2212 Rpbi , i = 1 . . . 4 (3) where the unit vectors of cable are: l\u0302i = li \u2016li\u2016 (4) The existence of pulleys will affect the inverse kinematics of cables as expressed inEq. (3), since therewill be additional cable length from point Ai to contact pointCi . Let the pulley center of radius r be denoted by point Mi . The coordinates of point M expressed in the pulley frame Fai is defined as: amT i = [ r 0 0 ] . To express the pulley center with respect to the reference frame, rotation matrices Rz and Ry are introduced to transform them to the base frame coordinates Fb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000006_ls.1481-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000006_ls.1481-Figure7-1.png",
        "caption": "FIGURE 7 The effect of inclined angle and included angle on load\u2010carrying capacity (LCC) and efficiency parameter [Colour figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            " By taking computation time and accuracy into consideration, radial grid number is selected as 170 in the following analysis. As mentioned above, the orientation, shape, and distribution of dimples on face seal are mainly decided by several geometrical parameters: long axis ratio A, short axis ratio B, inclined angle \u03b8, long axis included angle \u03b1, unit number n, and radial dimple number Nr. The influence of inclined angle \u03b8 and long axis included angle \u03b1 on LCC and efficiency parameter of gas face seals when unit number n = 2 are presented in Figure 7. It can be seen that both LCC and efficiency parameter are mainly effected by inclined angle, rather than long axis included angle, for the flow direction of sealing gas in the dimple is mainly determined by inclined angle. Both LCC and efficiency parameter increase or decrease periodically with the increases of inclined angle, especially for the conventional ellipse dimple when \u03b1 = 180\u00b0. Two excellent performance regions appeared for large LCC or large efficiency parameter, respectively, under different values of inclined angle and long axis included angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002459_iciea.2016.7603796-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002459_iciea.2016.7603796-Figure2-1.png",
        "caption": "Fig. 2. The film thickness at a point",
        "texts": [
            " ( ) 1 cos (2 , 2 ) 2 3cos (2 ,2 ) 3 2 t t k k y t t k k \u03c0\u03b4 \u03c9 \u03c9 \u03c0 \u03c0 \u03c9 \u03c0\u03b4 \u03c9 \u03c0 \u03c0 \u2208 \u2212 = \u2208 + (17) The gap between the moving ring and the static ring of the mechanical seal is 1 2h y y= \u2212 (18) If 2 20mm, 1 0.017r \u03b2= = = , then =0.34mm\u03b4 . Under normal circumstances, the thickness of the liquid film is only about 1 micron. And the leakage rate of mechanical seal is proportional to the three time of the liquid film thickness. Therefore, that the mechanical seal follows the poor has greatly increased the mechanical seal leakage rate. Effects of the following characteristics to contact gap under the condition of high speed is shown in Fig.2. As shown in Fig.1, the effect of the following characteristics on the film thickness is relatively significant for the different periodic harmonic motion of the static ring. However, the film thickness increases with the increment of the ratio of the rotation rate and periodic harmonic motion of the static ring. And the film thickness at one point fluctuates with time as Fig. 3. 1362 2016 IEEE 11th Conference on Industrial Electronics and Applications (ICIEA) Influence of the ratio of the rotation rate and periodic harmonic motion of the static ring on the average of film thickness is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure2-1.png",
        "caption": "Figure 2. An oblique circular torus obtained as a surface trajectory in a RR dyad.",
        "texts": [
            " In the noteworthy paper, Fichter and Hunt13 proved that a plane that is bitangent to a general circular toroid intersects the surface along a couple of congruent circles. The pertinent proof is established by using the complex projective completion of the Euclidean space. In what follows, one will not consider imaginary geometric entities but will focus on the special case with a zero offset. When the offset (d) is zero and two R axes are not twisted by a right angle (j j 6\u00bc 90 ), the circular toroid becomes an OCT depicted in Figure 2. The case (j j \u00bc 90 ) corresponds to the right circular torus that is the standard torus. An OCT is produced by revolving an inclined circle (circular generatrix) whose plane contains the common perpendicular of the circle axis and the revolution axis. It is proved that, generally, the revolution of other circular generatrices lying on planes containing the OCT center produces the same OCT. Because of the obvious symmetry of the revolute surface with respect to any plane containing the revolution axis, a second generatrix is simply the plane-symmetric of the first one. In the case (j j \u00bc 90 ), the foregoing two circular generatrices are meridians and the other two circles are those of Villarceau\u2019s theorem. Refer to Figure 2; (O, i, j, k) is a Cartesian orthonormal frame of reference whose origin O is the intersection of the first R axis and the common perpendicular of the two R axes. The unit vector k is parallel to the first R axis, which is characterized by (O, k). The unit vector i is parallel to the common perpendicular of the two R axes. The point B is the intersection of the common perpendicular and the second R axis. The vector (OB)\u00bcB O is equal to bi. The number b, b> 0, is the length (or Euclidean norm) jjOBjj of the bar (OB)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001724_iros.2015.7353938-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001724_iros.2015.7353938-Figure1-1.png",
        "caption": "Fig. 1. Examples of balancing on non-rigid contacts during whole-body task execution.",
        "texts": [
            " The problem becomes more complex when the rigidity of the object in contact is unknown a priori to robotic controllers, which is usually the case in many scenarios. This paper aims at adapting whole-body motions of humanoid robots to unknown rigidity properties of the environment. This work is dedicated to whole-body balancing, and more generally whole-body control, with non-rigid, unilateral, frictional support contacts, for example, standing on a soft ground, or pushing against a compliant support contact with one hand while reaching for an object far away with the other hand (see Fig.1). The problems of the manipulation of compliant objects and the handling of unexpected disturbance forces are beyond the scope of this paper. Moreover, the proposed control approach does The authors are with -Sorbonne Universite\u0301, UPMC Univ Paris 06, UMR 7222, Institut des Syste\u0300mes Intelligents et de Robotique, F-75005, Paris, France -CNRS Centre National de la Recherche Scientifique, UMR 7222, Institut des Syste\u0300mes Intelligents et de Robotique, F-75005, Paris, France {liu, padois}@isir.upmc.fr not handle anticipatory aspects of balance, but it provides a reactive mechanism to maintain balance while multiple motion and contact tasks are being performed in a compliant environment",
            " Objective (2c) is a regulation term that minimizes the norms of the variables \u03c4 and q\u0308. This objective is useful for ensuring the uniqueness of the solution for redundant robots. As this objective may increase the errors of other elementary tasks, its weight Qr is set to a very low value compared to other objective weights. The whole reactive whole-body control system is summarized in Fig. 3. In this work, a multi-contact state is supposed to be in static equilibrium if there is no motion at support contacts. Take the reaching scenario shown in Fig. 1 for example, if the support hand pushes too slightly against the table surface, the reaction force generated by the table is too weak to support the robot\u2019s leaning posture, and the robot may not be able to provide enough joint torques to maintain its posture. In this case, the contact is not in equilibrium as the support hand will keep sinking with the table surface. If the hand pushes strongly enough, the contact equilibrium may be established. However, the multi-contact system usually does not have a unique equilibrium state due to the redundancy of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000594_icarsc.2015.38-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000594_icarsc.2015.38-Figure2-1.png",
        "caption": "Fig. 2: Tilting motions",
        "texts": [
            " This section introduces the tilt-quadrotor ALIV3 prototype and its concept. The ALIV3 dynamic model, including the model of the existing sensors and actuators, is also presented. The tilt-quadrotor ALIV3 platform (see fig. 1) consists of a structure with a central core and four arms, each with a motorpropeller set. Two opposing arms have fixed motor-propeller sets, while the other two arms have swivelled motor-propeller sets. The tilting motions are the result of the action of four servos. The two servos located in the swivel arms control the roll-tilt (\u03c6ti ) - see fig. 2a. The two other servos, located in the central core, are responsible for the pitch-tilt (\u03b8ti ) of the rotors (fig. 2b). Besides the pitch-tilt servos, the central core also allocates the battery, an ArduPilot APM1 (Arduino-based processor for flight control, including 3-axis accelerometer, gyroscope and 978-1-4673-6991-6/15 $31.00 \u00a9 2015 IEEE DOI 10.1109/ICARSC.2015.38 156 magnetometer, and a barometer), the power distribution board, and the electronic speed controllers (ESCs). The landing gear is also attached to the central core. The tilt-quadrotor differs from a standard quadrotor by the substitution of two normal fixed rotors for two tilting ones, resulting in the advantage of maintaining the payload leveled, independently of its motion or velocity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002571_2168-9806.1000135-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002571_2168-9806.1000135-Figure2-1.png",
        "caption": "Figure 2: Crawler track assembly interacting with the ground.",
        "texts": [
            " Their study developed the 3-D virtual prototype simulator of crawler track interacting with oil sands terrain within MSC ADAMS and reported the simulation results for crawler track shoe kinematic quantities (linear and angular displacement, velocity and accelerations). In this paper the time varying contact forces that are used as an input to calculate crawler shoe kinematic quantities reported in Frimpong and Thiruvengadam [12,13] are presented. Rigid multi-body dynamics of crawler-terrain interactions Figure 2 shows the geometry of crawler track assembly interacting with oil sand terrain. The track is modeled using the crawler track dimensions for the P&H 4100C Boss in Table 1. The crawler track is assembled from crawler shoes connected together by two link pin joints. One link pin is made a spherical joint and the other link pin is made a parallel primitive joint to create equivalent revolute joint between two crawler shoes. The crawler shoe model is generated in Solidworks based on the actual crawler shoe model for P&H 4100C Boss shovel [14]",
            " Similarly, the oil sand terrain is made up of spring- damper-oil sand units connected to four adjacent oil sands units by spherical joints [10]. The stiffness (k) and damping (c) values of oil sand terrain are listed in Table 2. Only the open track chain of the crawler assembly, in contact with the ground Figure 1, is used for this study. More details on the dimensions, joints and material properties of crawler shoe and oil sand units can be found in Frimpong and Thiruvengadam [12,13]. The global coordinate system is located at the left corner of the oil sand terrain at point O as shown in Figure 2. The position of the center of mass of crawler shoes 1 - 13 and oil sand units 15 and 64 at time t = 0 with respect to the global coordinate system are listed in Table 3. The joint locations of spherical and parallel primitive joints between each crawler shoes are listed in Table 4. Governing equations of motion and solution methodology The kinematic and dynamic equations of motion that govern the propelling motion of the rigid crawler track on oil sand terrain based on multi-body dynamics theory [15,16] can be found in Frimpong and Thiruvengadam [12,13]",
            " The external forces acting on the crawler shoe # i are the gravity force (mig) due to self-weight of the shoe, uniformly distributed load (wi) due to machine weight and contact forces ,i i c cF T due to interaction between crawler shoe i and ground as shown in Figure 3. The joint forces at the spherical joints ( 1, 1,,\u2212 \u2212i i i i s sF T and , 1 , 1,+ +i i i i s sF T ) and parallel primitive joints ( 1, 1,,\u2212 \u2212i i i i p pF T and , 1 , 1,+ +i i i i p pF T ) are also shown in Figure 3 For 63 interconnected rigid multi-body system shown in Figure 2, the differential-algebraic equations of motion can be written from Shabana and MSC [16,18] as in equation (1). + = + T u a gMu K Q Q\u03bb ( , ) 0=tK u (1) M - Mass matrix of the system; u - vector of system generalized coordinates; K - constraint equations due to joints and applied motion; Qa - Applied forces; Qg - gyroscopic terms of the inertia forces. Frimpong and Thiruvengadam [13] presented detailed formulation for the generalized inertia and external forces acting on the crawler shoes and oil sand unit",
            " In addition, the crawler shoe joint constraint forces and torques and total deformation of the oil sand terrain are presented Volume 4 \u2022 Issue 2 \u2022 1000135 J Powder Metall Min ISSN: 2168-9806 JPMM, an open access journal as part of the solution to the equations of motion. The normal and frictional contact force results play a major role in the dynamic stress analysis and subsequent fatigue study of the flexible crawler shoes. The reaction forces and torque acting at the crawler shoe joints represent the loading at the pin joint and has a direct bearing on the link pin wear and tear which in turn will impact the crawler shoe fatigue life [2]. The virtual prototype of the crawler track assembly shown in Figure 2 is modeled in MSC ADAMS [12] to simulate the crawler propel action on oil sand terrain for the two types of motion constraints. The virtual simulation is carried out for the time period of 10 s and the resulting dynamic variation of contact forces on each crawler shoes, reaction forces from the crawler shoe joints and total deformation of the oil sand terrain are presented in this section for the two motion constraints. Case 1: Only translation: The comparison of the time distribution of crawler shoe sinkage into the oil sand and sinkage velocity at the contact point is shown in Figures 5a and b for Parts 2, 5, 8, 11 and 14"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000155_ecce.2019.8913088-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000155_ecce.2019.8913088-Figure1-1.png",
        "caption": "Fig. 1. DT-SynRM with W-12-12.",
        "texts": [
            " Single-SSF and triple-SSF cause a large short circuit current. For this reason, the open-circuit and short-circuit faults result in high electromagnetic torque ripple and serious mechanical vibration [6]. Therefore, the post fault analysis is necessary to predict the fault condition and prevent further problems or accidents. In [7], three different winding arrangements are described such as two in-phase models (W-11-22 and W-12-12) and one out-of-phase model (W-6phase). This study just focuses on only one winding arrangement of W-12-12, as shown in Fig. 1. First, inductance behaviors are analyzed by Finite element method (FEM) and they are verified by measurement of induced voltages. Next, torque and current characteristics are analyzed under healthy, half control (HC)-mode and different fault conditions. The fault conditions are analyzed under single-OPF, single-OSF, single-SSF, and triple-SSF. DT-SynRM with two sets of winding is considered to investigate the characteristics under healthy, HC-mode and different fault conditions. The DT-systems are refered to as ABC-phase system and XYZ-phase system, which is supplied 978-1-7281-0395-2/19/$31"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000847_s10556-013-9774-9-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000847_s10556-013-9774-9-Figure1-1.png",
        "caption": "Fig. 1. Diagram of displacement pump compressor with gas damper (the basic thermodynamic parameters of the control volumes and circulating flows of fluid and gas are indicated).",
        "texts": [
            " When a pump and compressor are combined into a unified assembly, a pump with high angular spin rate of the drive shaft is required, since the values of the optimal rotational speeds of the compressor are substantially higher. In this case, one of the simple alternatives is to create a displacement pump with gas damper, whence one portion of the fluid will be fed to the customer in the course of discharge while the other portion of the fluid will be delivered to a gas cavity fabricated into a piston (Fig. 1). By decreasing the delivery of the pump, it becomes possible to reduce losses of pressure in the course of discharge and increase the operating efficiency of the pump. A calculation of processes of compression and expansion in a displacement pump with gas damper is presented in [1]. The objective of the present work is to develop methods of calculating processes of discharge and suction in a dis- placement pump with gas damper. In performing theoretical studies, we will adopt assumptions analogous to those of [1], i.e, a dropping liquid is com- pressible and obeys Hooke\u2019s law; a compressible fluid obeys the law of an ideal gas; the distribution of the thermodynamic parameters in the gas and liquid cavities of the pump is uniform; the walls of the working cavity of the pump are absolutely rigid; dissolution of gas in the liquid is negligibly low; the interface in the gas plane is a surface level and disturbance of the continuity of the liquid is absent. Calculation of the Discharge Process (Fig. 1). In the course of discharge, a portion of the liquid will be displaced through the discharge valve to the customer (dM15), and a portion will enter the gas cavity (dM17), moreover leakage of liquid will be observed, through leaks in the suction valve (dM14), the displacement seal (dM12), and through the rod seal (dM19). In this case, we may write the base equation for preservation of volume in the form (1)dV dM dM dM dM dV dVw wk g\u2212 + + +( ) = +15 12 14 19 / ,\u03c1 Chemical and Petroleum Engineering, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003329_aim43001.2020.9158972-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003329_aim43001.2020.9158972-Figure11-1.png",
        "caption": "Figure 11. Experimental setup",
        "texts": [
            " Next, the robot is actually crawled in the pipeline, and the sensor values are acquired. From the acquired sensor values, the crawling distance and transit time between the straight pipe section and bending pipe section, and bending pipe that passed through are obtained. By combining these, the pipeline can be restored. In this section, in order to evaluate the proposed distance measurement method, the moving distance of the IMU sensor attached to the guide rail was calculated during movement. Fig. 11 shows the experimental environment. An IMU sensor was mounted on a guide rail that moved in a single-axis direction. This time, for simplicity, the sensor was installed so that the moving direction of the sensor was in the positive xaxis direction. In addition, the data obtained from the sensor was transferred to the PC via a USB connection. The sampling time was set to 5 ms. The IMU sensors used were the same as those mounted on the robot. The IMU sensor attached to the guide rail was set to motion to obtain the value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001756_978-3-319-24502-7_10-Figure10.4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001756_978-3-319-24502-7_10-Figure10.4-1.png",
        "caption": "Fig. 10.4 A cube translating through a viscous fluid with velocity V under the influence of force F acting on its centre. For highly symmetric particles, linearity of Stokes equation implies that the force and velocity are collinear, with the drag force being independent of the particle orientation",
        "texts": [
            " The particle experiences then the drag force \u2013 F which we can decompose into forces acting along the axes of the coordinate system according to F = {Fi}. From linearity, we conclude that the force F1 acting on a particle moving with velocity (V1,0,0) must be of the form F1 \u00bc aV1, with a being a positive constant. Imagine now that the particle is a cube with its edges aligned along the coordinate axes. Then, from symmetry, F2 \u00bc aV2 and F3 \u00bc aV3, and in general F \u00bc aV. Hence the drag force experienced by a cube does not depend on its orientation, and it is collinear with the velocity (see Fig. 10.4). As everyday experience teaches us, this is obviously not valid any more for large Re. In fact, a more general statement is true in Stokes dynamics: Any homogeneous body with three orthogonal planes of symmetry (such as spheroids, rods, cylinders, disks, or rings), will translate under the action of force without rotating, although in general with a sidewise velocity component perpendicular to the driving force. The sidewise motion is absent only if the force is acting along the rotational symmetry axis of the particle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000882_0954405414564405-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000882_0954405414564405-Figure10-1.png",
        "caption": "Figure 10. Axial hobbing method.",
        "texts": [],
        "surrounding_texts": [
            "Pressure angle a ( ) 20 Pressure angle a ( ) 20 Number of teeth ZC 15 Number of hob threads ZB 1 Normal module mn (mm) 6 Normal module mn (mm) 6 Total tooth depth h (mm) 6 Hob spiral angle n ( ) 10 Tooth width B (mm) 30 Spiral angle direction Right hand Outside diameter of gear blank DC (mm) 120 Outside diameter of hob DB (mm) 20 Gear spiral angle b ( ) 30 Spiral angle direction Left hand Installation height H (mm) 20 at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from workpiece. Then, the EGB was opened by the G81 command, and the hob was gradually fed to the full tooth depth along the X-axis. After a short pause, the hob was fed along the Z-axis over the entire tooth width, as the hob cut along the X-axis and Z-axis. Finally, the EGB was closed by the G80 command. During the cutting process, the single-axis tracking error of the X-axis and Z-axis was both less than 0.4mm, and the tracking error of the C-axis was less than 0.0006 rad. The pitch error of the workpiece reflects the accuracy of EGB and is illustrated in Figure 15, with a maximum absolute error less than 15mm. Diagonal hobbing method. In this section, the diagonal hobbing method is evaluated. In this case, the spindle speed was 900 r/min. The reference and actual trajectories are illustrated in Figure 16. The C-axis followed the B-axis, Y-axis and Z-axis throughout the machining process, and its reference position and actual position are illustrated in Figure 17. There is also an instantaneous position change at the moment when the EGB is opened, as the C-axis attempts to keep up with the spindle. The tracking error during the diagonal hobbing motion process is shown in Figure 18. It can be seen that the single-axis tracking errors of the X-axis, Y-axis and Z-axis are all less than 0.4mm during the hobbing process, and the tracking error of the C-axis is less than 0.0006 rad. The estimate of the gear pitch error in the at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from gear diagonal hobbing process is illustrated in Figure 19, and the maximum absolute error is less than 18mm. From Figures 12\u201319, we can conclude that the hobbing CNC and the EGB can work correctly for the examples tested. The C-axis follows the rotation speed of the spindle and the feed rate of the other servo shaft. From Figures 14 and 18, we can see that the tracking error of each axis is very small. Here, we use the gear pitch error as the performance metric of the EGB. As shown in Figures 15 and 19, under the control of the EGB, the maximum gear pitch error is 0.0147mm in the axial hobbing movement process, and the maximum at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from gear pitch error is 0.0176mm in the diagonal hobbing movement process. The results show that the proposed EGB is effective."
        ]
    },
    {
        "image_filename": "designv11_30_0003109_j.matpr.2020.04.299-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003109_j.matpr.2020.04.299-Figure1-1.png",
        "caption": "Fig 1: Two-lobe fluid-film bearing.",
        "texts": [
            " In current study the effect of variation in couple stress parameter for a given concentration of nanoparticle additives on static and dynamic performance parameters of two-lobe journal bearing working with nano-lubricant suspended with TiO2 nanoparticle. At a fixed couple-stress parameter pressure profile is plotted for different concentration of nanoparticle. Krieger-Dougherty Model for viscosity is used to notice viscosity variation. Reynolds equation is modified for couple-stress parameter and same is solved by Gauss-Siedel method with the help of successive over relaxation technique in a finite difference grid. 2. Theory: Schematic diagram of two-lobe journal bearing is shown in Fig. 1. Lobe 1 and 2 has its centre at 1O and 2O respectively. Lobes are circular in geometry separately but bearing as a whole is non-circular. A groove of size 100 is provided to both the lobes for the purpose of oil supply. Center to center distance between individual lobe and bearing is known as ellipticity ( pe ) and ellipticity ratio is defined as pe C\u03b4 = . Current study considers ellipticity ratio as 0.5. In this study relative viscosity is regarded as a function of concentration of solid nanoparticle in the base lubricant and Krieger-Dougherty Model is used to obtain relative viscosity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002632_tencon.2016.7848445-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002632_tencon.2016.7848445-Figure6-1.png",
        "caption": "Fig. 6: Selective laser sintered parts",
        "texts": [
            " The SLS process takes the STL code and builds the parts inside a powder bed with the moveable platform. After the printing process is finished, the powder bed is removed from the machine and printed part is separated using high velocity jet air. Then the remaining powder will reused for the subsequent printing process. SLS printed parts are the stronger than the SL parts as the materials are fused at high temperature and treated after the printing process. Parts that require good mechanical strength like turbine blades, columns can be printed as shown in Fig. 6. SLS provide good mechanical strength as it uses metal powders like Aluminum, Titanium, alloys etc. 2328 2016 IEEE Region 10 Conference (TENCON) \u2014 Proceedings of the International Conference Fused deposition modelling[3][7] is simple technique which utilizes thermoplastic filaments as the source of material. (shown in Fig. 7). Same as all the 3D printers, FDM machine also requires the STL files as the input. The machines has a printer head wherein the thermoplastic filament is held at a nozzle at temperature higher than the melting point of the material at the tip of the nozzle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003778_ssrr50563.2020.9292574-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003778_ssrr50563.2020.9292574-Figure9-1.png",
        "caption": "Fig. 9. Simulation scenario consisting of a pallet, rough terrain, and a ramp.",
        "texts": [
            " We obtain the angle as the minimum over all cells \u03b1\u0303 = atan2 (hground, dheightmap) Please note that this simplification has a small systematic error as indicated by the dashed line in Fig. 6. Since the rotation axis in effectively all cases does not pass through the origin, the origin of the robot has to be moved as well to clear the rotation-induced translation. The accuracy of the introduced method for predicting the robot pose was evaluated on the real robot in the scenarios depicted in Fig. 7 and the stability predictions were evaluated in a simulated scenario depicted in Fig. 9. a) Real Robot Scenarios: The ramp scenario was repeated for three inclinations at 5\u00b0, 12\u00b0, and 20\u00b0. For each inclination, the robot was rotated 360\u00b0 in steps of 22.5\u00b0 and the pose was recorded. In the elevated ramp scenario, the robot was driven across the elevated ramps manually while using the reconfigurable flippers to test how the approach handles complex poses where the robot is only supported by its flippers. The robot was stopped several times along the path and the pose and joint states were recorded to minimize the influence of dynamic effects on the pose",
            " For example, using a contact threshold of 0.5 cm for the first iteration and rotation estimation and a threshold of 2.0 cm for the stopping criterion, the mean error is at 1.54\u00b0 \u00b11.08\u00b0. b) Simulated Scenario: As capabilities to obtain ground truth contact points to test the stability estimates of the presented approach were not available, it was evaluated with the physics simulation engine Open Dynamics Engine (ODE) used in the robotics simulator Gazebo 9 as a reference. The simulation scenario depicted in Fig. 9 features a step, rough terrain made of boxes of varying height, and a ramp. The path of the robot straight across the set of obstacles was simulated and the poses, contact points, and a ground truth heightmap of the scenario were obtained using the open-source package hector ground truth gazebo plugins 2 developed in the context of this work. Fig. 10 shows the predicted and simulated roll and pitch, and the stability calculated using the FASM for a map 2https://github.com/tu-darmstadt-ros-pkg/hector_ ground_truth_gazebo_plugins resolution of 5 cm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000594_icarsc.2015.38-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000594_icarsc.2015.38-Figure4-1.png",
        "caption": "Fig. 4: Tilt-quadrotor propulsion forces and moments, and inertial and body-fixed reference frames",
        "texts": [
            " \u2022 The forward x-axis motion requires rotors 2 and 4 to have pitch-tilt angles with opposing signals and a higher thrust than rotors 1 and 3. \u2022 The lateral y-axis motion requires rotors 2 and 4 to have roll-tilt angles with opposing signals and a higher thrust than rotors 1 and 3. \u2022 The yaw motion is obtained by tilting rotors 2 and 4 with pitch-tilt angles with the same signal. A yaw motion is not possible without tilting the rotors since it results in a null z-moment. Consider the reference frames represented in fig. 4, where I = {OI ;xI , yI , zI} is a fixed (inertial) frame assumed here as the north-east-down (NED) frame, and B = {OB;xB, yB, zB} is the body-fixed frame. The rotation matrix R \u2208 SO(3) = {R \u2208 R 3\u00d73|RTR = I, det(R) = 1} represents the orientation of the body frame B with respect to the inertial frame I. Using the Euler angles \u03a6 = [\u03c6, \u03b8, \u03c8]T attitude description, R is given by R = ( c\u03c8c\u03b8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03c6 c\u03c8s\u03b8c\u03c6 + s\u03c8s\u03c6 s\u03c8c\u03b8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03c6 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 ) (1) where the short notation c\u00b7 = cos(\u00b7) and s\u00b7 = sin(\u00b7) was used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003538_00405000.2020.1827580-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003538_00405000.2020.1827580-Figure2-1.png",
        "caption": "Figure 2. The structure of helical yarn.",
        "texts": [
            " As the flexible property compared with the wrap filament, only the effect of the normal force on the cross section and radial strain of the core filament was mainly studied and analyzed. 5. The auxetic complex yarn was with uniform and stable structure. And one helical element could be selected to represent the whole yarn for the theoretical analysis. To analyze the effect of interaction force on the cross section and radial strain of the core filament, mechanical model was selected as shown in Figure 2. With the fixed Oxyz coordinate system, the initial helical angle of a single filament was ac. The helical core filament was applied of axial tension Tc and radial normal force, which were related to F1, F2 and N in Figure 2. Besides, the P-uvw coordinate system with any point P on the helix was as the moving coordinate system. Firstly, F1 and F2 could be calculated by Tc and ac as follows: F1 \u00bc Tccosac (2) F2 \u00bc Tcsinac (3) According to Batra (1973), the normal force N on the contact surface was decided by the axial tension Tc, helical angle ac and radius Rc, as N \u00bc Tc sin2ac Rc (4) If\uff08exn, e y n, ezn\uff09meant the direction vectors, the moments of N in three directions could be expressed as follows: MNX \u00bc \u00f0s 0 NeynZds\u00fe \u00f0s 0 NeznYds (5) MNY \u00bc \u00f0s 0 NexnZds\u00fe \u00f0s 0 NeznXds (6) MNZ \u00bc \u00f0s 0 NexnYds\u00fe \u00f0s 0 NeynXds (7) Besides, the bending and twisting moments G, G\u2019 and H of helical filament were acquired from reference Love (Love, 1944) G \u00bc B\u00f0p p0\u00de (8) G0 \u00bc B\u00f0q q0\u00de (9) H \u00bc C\u00f0r r0\u00de (10) where B and C were flexural and torsional rigidities; p0, q0, r0, p, q and r were initial and deformed curvature and twist"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002390_gt2016-57410-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002390_gt2016-57410-Figure3-1.png",
        "caption": "Figure 3. Front bearing configuration",
        "texts": [
            " In this way, the seals before the impulse wheel are characterized by high cross-coupling stiffness coefficients. 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89517/ on 03/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use For the initial configuration, the turbine rotor, shown in Figure 2, was supported by two tilting pad journal bearings and a double tilting pad thrust bearing; 3D models and cross sections of the bearings are shown in Figure 3 and Figure 4. The shaft sizes and geometric characteristics of the journal bearings are listed in Table 1. The front bearing, located at the steam admission side of the machine, was a horizontally split journal bearing combined with a selfequalized double thrust bearing with direct lubrication, arranged in a back to back configuration between the two thrust collars. The tilting pad thrust bearings consisted of 6 offset pivot steel backed pads on each thrust face, supplied with oil using bi-directional spray nozzles between the thrust pads"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002115_cjme.2015.1015.122-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002115_cjme.2015.1015.122-Figure1-1.png",
        "caption": "Fig. 1. General 6-6 Stewart mechanism",
        "texts": [
            " Bivectors 32e , 13e and 21e correspond to the quaternion basis i , j and k . Translation in 3 1,1+ can be defined as 1 1 , 2 1 1 , 2 e e \u00a5 * \u00a5 = - = + T a T a (5) where 2d=a t in which d denotes the displacement distance and t denotes the direction of movement. Therefore, the motion transformation of a rigid body in 3 1,1+ can be denoted by .\u00a2 = * *Q TRQR T (6) WEI Feng, et al: Algebraic Solution for the Forward Displacement Analysis of the General 6-6 Stewart Mechanism \u00b758\u00b7 2.2 Modeling using CGA A kinematic model of a general 6-6 Stewart mechanism is shown in Fig. 1. The six leg lengths provided by the prismatic joints in each leg are the six inputs to control the location and orientation of the moving platform. For the moving and fixed platforms, the respective spherical joints iB and iA ( 1, 2, ,6)i =  are not restricted to lie in a plane. We locate a fixed frame 1 1 1 1O x y z for the fixed platform with its origin at 1A , and attach a moving frame 2 2 2 2O x y z to the moving platform with its origin at 1B . Let T( , , ) ( 1, 2, ,6)i xi yi zia a a i =a  denote the coordinate of the fixed spherical pair iA in the fixed frame, where the transpose is represented by T, T( , , )i xi yi zib b bb ( 1, 2, ,6)i =  denote the coordinate of the spherical pair iB in the moving frame, iL denote the ith leg length, T( , , )x y zp p pp denote the translation vector from 1A to 1B in the fixed frame, and the rotational matrix R denote the orientation matrix of the moving platform with respect to the fixed platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003270_s42243-020-00467-0-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003270_s42243-020-00467-0-Figure1-1.png",
        "caption": "Fig. 1 Electronic additive manufacturing and steel part. a Experimental setup of EAM; b EAM 16MND5 steel; c schematic representation of tested steel in vertical view",
        "texts": [
            " It is expected that the impact toughness of EAM steel is similar to or better than the standard level of CF steel. CF 16MND5 steel was molten in a vacuum induction furnace and casted into an ingot. The ingot was forged into annular tested steel in a temperature range of 900\u20131150\u00a0\u00b0C, with diameter of 1210\u00a0mm, and wall thickness and height of 150\u00a0 mm. Subsequently, the annular tested steel was quenched after austenitizing at 900\u00a0\u00b0C for 5\u00a0h and then tempering at 650\u00a0\u00b0C for 4\u00a0h. The experimental setup of EAM is shown in Fig.\u00a01a. The wire was molten and deposited on the substrate by a heating source. EAM 16MND5 steel was prepared by layer-by-layer materials melting and depositing. The motion of heating source and wire feed nozzle was controlled by a numerical control program. Then, EAM 16MND5 steel was tempered at 650\u00a0\u00b0C for 4\u00a0h to eliminate residual stress and machined to remove iron oxide scale, as shown in Fig.\u00a01b. Because the impact toughness of CF 16MND5 steel was affected by the wall thickness, EAM 16MND5 steel is manufactured to have the same geometry and size. In addition, the two manufacturing processes were completely different, and in order to obtain their own best impact toughness, each adaptive chemical composition was selected. The chemical composition of the tested steel is shown in Table\u00a01, which meets the standard of RCC-M specification. The contents of C and Cr were reduced, and the content of Si was increased, 1 3 which is beneficial to prevent EAM steel from cracking during the manufacturing process",
            " It is worth noting that the direction of the specimens in the tensile and impact tests was vertical to that of deposition, while the section of the specimen in the hardness tests and microstructure observations was parallel to that of deposition. During service, RPVs were subjected to fluid flushing which is similar to an outward thrust. Therefore, the direction of the impact specimen should be perpendicular to this thrust. At the same time, RPVs were operated under high pressure of 17\u00a0MPa, and the wall thickness was subject to tensile stress along the circumferential direction. The geometry and direction of tested steel are shown in Fig.\u00a01b, c. After testing, the mechanical properties of EAM specimen are consistent at all wall thicknesses, so that all test specimens in this article were sampled at 1/2 wall thickness. OM and SEM investigations of CF specimen are illustrated in Fig.\u00a02. As shown in Fig.\u00a02a, the microstructure of CF specimen is typical bainite with a large number of carbides [19]. It can be clearly seen in Fig.\u00a02b that coarse carbides are distributed along the grain boundaries and in the ferrite matrix of bainite, respectively",
            " Wp1 value of CF specimen is higher than that of EAM specimen, while Wp2 and Wp3 values are just the opposite. For EAM specimen, less energy is consumed when microvoids combine to form macrocracks, which is responsible for the lower Wp1 value. The reason will be discussed in Sect.\u00a04.4. Since multiple HAGBs in EAM specimen cause the twist of crack propagation path to consume more energy, Wp2 and Wp3 values of EAM specimen are higher than that of CF specimen. The conclusion is also identified by the fracture morphology, as shown in Fig.\u00a01b, c, e, f. The fibrous region, radial region and shear lip represent the stable crack propagation, unstable crack propagation and post-unstable crack propagation phases in the impact curves, respectively. Usually, the fibrous region and shear lip indicate more energy consumed than radial region. Comparing Fig.\u00a012b, e, CF specimen has larger fibrous and radial regions than those of EAM specimen, while the shear lip does the opposite. This indicates that Wp2 and Wp3 values of EAM specimen are higher than CF specimen, while Wp1 value is reversed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001150_00207721.2015.1018367-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001150_00207721.2015.1018367-Figure1-1.png",
        "caption": "Figure 1. Two inverted pendulums connected by a spring.",
        "texts": [
            " While using Theorem 1, a novel approach is proposed to study the practical output tracking problem for switched nonlinear systems in p-normal form. 4. Examples In this section, we show the applicability and effectiveness of our approach on two examples illustrating the main results of the paper. Example 1: Let us demonstrate our constructive practical output tracking scheme via a mechanical system as that has been handled in Spooner and Passino (1999), i.e., we consider the control of two inverted pendulums connected by a spring as depicted in Figure 1. Each pendulum may be positioned by a torque input ui applied by a servomotor at its base. The equations of motion for the pendulums are described by x\u03071 = x2, (29a) x\u03072 = ( m1gr J1 \u2212 kr2 4J1 ) sin x1 + kr 2J1 (l \u2212 b) + u1 J1 + kr2 4J1 sin x3, (29b) x\u03073 = x4, (29c) x\u03074 = ( m2gr J2 \u2212 kr2 4J2 ) sin x3 \u2212 kr 2J2 (l \u2212 b) +u2 J2 + kr2 4J2 sin x2, (29d) y = x1, (29e) where (x1, x3)T = (q1, q2)T and (x2, x4)T = (q\u03071, q\u03072)T are the angular displacements of the pendulums from vertical and angular rates, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000256_icar46387.2019.8981551-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000256_icar46387.2019.8981551-Figure2-1.png",
        "caption": "Fig. 2. (a) Photograph and (b) antenna indices of the 50-element antenna array used in the Tx and Rx of the channel sounder.",
        "texts": [
            " EMPIRICAL STOCHASTIC POLARIZATION MODELS FOR INDOOR SCENARIOS In this section, the modeling approach proposed here is applied to establish some empirical stochastic propagation models based on wideband channel measurements. Different statistics of the ellipse parameters observed from indoor scenarios of various types demonstrate the necessity of refined modeling of the polarizations using the proposed approach. The measurements were performed by Elektrobit Oy and Technology University of Vienna in 2005 using the wideband MIMO channel sounder\u2014PROPSound\u2014in a building at Oulu University. A 50-element and a 32-element patch-antenna arrays were used in the Tx and Rx, respectively. The photograph illustrated in Fig. 2(a) depicts the appearance of the 50-element array. Fig. 2(b) shows the indices of antennas on the array. The 32-element Rx array has the same configuration as the 50-element Tx array. However, the antennas from No. 19\u201336 were not activated for receiving signals during the measurement. Each patch in the array consists of two antennas which are +45\u25e6 and \u221245\u25e6 polarized, respectively. The measurements were conducted with effective bandwidth of 100 MHz at the carrier frequency of 5.25 GHz. The sounding of individual subchannels between any pair of Tx and Rx antennas is performed sequentially in a time-division-multiplexing (TDM) mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure3-1.png",
        "caption": "Fig. 3. Mesh of the recirculating hollow-shaft cooling structure.",
        "texts": [
            " In the case of steady-state operation, the RANS equations are as follows: 0i i V x \u2202 = \u2202 (1) ( ) ( )2i j ij i j j i j P VV S v v x x x \u03c1 \u03bc \u03c1\u2202 \u2202 \u2202 \u2032 \u2032= \u2212 + \u2212 \u2202 \u2202 \u2202 (2) where V, v are the fluid time-averaged and fluctuating velocity component; \u03bc is the fluid dynamic viscosity; \u03c1 is the fluid density; pC is the specific heat; ijS is the mean strain-rate tensor, i jv v\u03c1 \u2032 \u2032\u2212 is the symmetric Reynolds stress tensor. The solution accuracy is higher than that of finite elements analysis (FEA). At the same time, CFD has higher criteria for mesh quality, which requires more computing resources and is slow to solve. A three dimension of hollow shaft is modeled and its three-dimensional mesh is depicted in Fig. 3. The CHTC and friction loss are the focal points of this paper, so a model that can more accurately consider the effects of rotation and high shear stresses to solve the fluid domain is needed. The shear stress transport (SST) turbulence model was chosen since it can generally provide better performance than the Realizable k-epsilon model as it is more suitable for swirling flows without requiring sublayer damping [18] and includes a nonlinear cubic constitutive relation, which accounts for the anisotropy of turbulence [12]",
            " It is noted that friction loss of recirculating hollow-shaft is higher than direct-through hollow-shaft. III. STRUCTURE OPTIMIZATION OF RECIRCULATING HOLLOW-SHAFT From the analysis of the CHTC above, it is noted that the CHTC in the middle region is small and uniform compared to bottom wall and top wall of the recirculating hollow-shaft. In Section II, the phenomenon of a small CHTC at the bottom of the recirculating hollow shaft is analyzed when the rotational speed is 3000 rpm. In order to solve this problem, the bottom tail baffle shown in Fig. 3 is changed into a coneshape. Fig. 10 shows the velocity contour and vector diagram of the recirculating hollow-shaft with cone shape. Comparing Fig. 10 and Fig. 6(b), it can be seen that the cone angle facilitates the formation of a stronger toroidal vortex flow in the bottom area. As shown in Fig. 5(b), the CHTC at the bottom of the original structure is around 250 W/m2/K. However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000558_icra.2015.7139946-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000558_icra.2015.7139946-Figure3-1.png",
        "caption": "Fig. 3. Catheter Model Configuration Parameters, including insertion curvature \u03ba, arc length s, articulation angle \u03c6, and orientation angle \u03b2, and Control Tendon Inputs \u2206l1 through \u2206l4 and Insertion Axis Input \u2206l5",
        "texts": [
            " This testbed utilizes a trakSTAR system (Ascension Technologies, Burlington, VT) where the sensors instantly measure the transmitted field vectors and compute their real-time position and orientation relative to the transmitter. A sensor is placed through the center lumen of the prototype into the center of the distal cap and calibrated to the catheter. A mechanics based model is used for catheter control and localization, mapping manipulated inputs to the output catheter endpoint position. The model, depicted in Figure 3, is based on one originally presented in [8] and later expanded in [22] and [14]. This model assumes constant curvature and no friction between the control tendons and supporting lumens. Modifications, however, have been made to accommodate the insertion axis degree of freedom and to allow additional degrees of freedom in the modeling parameters. The state vector x represents the position coordinates of the catheter fiducial, corresponding to the endpoint fiducial frame {f}, in the reference frame {r} as x = ( X Y Z )> (1) where {r} is defined affixed to the end of the proximal sheave on the catheter fixture",
            " The development of the model, estimating the state output based on the manipulated tendon inputs, is detailed as follows. The prototype catheter is manipulated through the control actions of the CREST drive motors. These manipulated inputs are represented as \u03b8 = ( \u03b81 \u03b82 \u03b83 \u03b84 \u03b85 )> (2) where \u03b81 through \u03b84 are the angular articulation axes positions corresponding to the tendon control wires and \u03b85 is the angular position of the insertion axis lead screw. The tendon and insertion displacements depicted in Figure 3 are represented by the vector y = ( \u2206l1 \u2206l2 \u2206l3 \u2206l4 \u2206l5 )> (3) referred to as the catheter joint space. The relationship of motor inputs to joint space is given by y = Rs\u03b8 (4) where Rs is a diagonal gain matrix of articulation axes spool radii and the insertion axis lead screw pitch. Assuming the catheter to be quasi-static, the authors in [8] derive moment balance, about the x and z axes, and force balance equations yielding ( Kbx Kbz Ka ) qc =  T1d1 \u2212 T3d3 T4d4 \u2212 T2d2 4\u2211 i=1 Ti  (5) where qc is the configuration state vector",
            " Then, given a joint vector y and assuming \u03b5a 1, an iterative solution is formulated by constructing a linear system of the form Mq\u0302c = b (10) where M = l\u0302b(c1d1 \u2212 c3d3) l\u0302b(c4d4 \u2212 c2d2) Ka + la 4\u2211 i=1 ci Kbx + l\u0302b(c1d1 + c3d3) 0 la(c1d1 \u2212 c3d3) 0 Kbz + l\u0302b(c2d2 + c4d4) la(c4d4 \u2212 c2d2)  (11) and b = ( 4\u2211 i=1 ci\u2206li (c1d1\u2206l1 \u2212 c3d3\u2206l3) (c4d4\u2206l4 \u2212 c2d2\u2206l2) ) > (12) This iterative approach recursively estimates lb and qc until a convergence criterion is met defined as the iteration when the difference between successive values of \u03b5\u0302a fall below a tunable threshold value, \u03b4. For all experiments conducted, convergence occurred within 2 time steps. This algorithm, symbolized by A(y), is summarized in Table I. Function A(y)r Set \u03b5\u0302a = 0 and calculate l\u0302br Construct M, b, and solve q\u0302c = M\u22121br doa Set \u03b5a = \u03b5\u0302a and calculate l\u0302ba Construct M and solve q\u0302c = M\u22121br while |\u03b5a \u2212 \u03b5\u0302a| > \u03b4r if \u03b5\u0302a <= 0a Set \u03b5a = 0 and compute \u03bax and \u03bazr return qc = q\u0302c TABLE I JOINT TO CONFIGURATION FUNCTION, A(y), ALGORITHM In Figure 3, the catheter is shown in a general configuration described geometrically by the curvature \u03ba, the arc length s along the central axis, and the angle \u03b2 measured counter-clockwise from the x axis to the projection of the catheter onto the xz plane. These geometric parameters are defined in [23] as s \u03ba \u03b2  =  lb(1\u2212 \u03b5a) \u221a \u03bax 2+\u03baz 2 1\u2212\u03b5a arctan \u03bax \u03baz  (13) The coordinates of the catheter state vector are then derived in these geometric terms as X Y Z  =  cos \u03b2 \u03ba (1\u2212 cos s\u03ba) sin s\u03ba \u03ba \u2212 sin \u03b2 \u03ba (1\u2212 cos s\u03ba)  (14) Equations 13 and 14 are encapsulated in a configuration to joint space function symbolized by H(qc)",
            " Developing a robust control system which operates the catheter through trajectories similar to surgical procedures provides a framework for development of increasingly autonomous robotic surgery. The control system, depicted in Figure 4, utilizes a PI controller with State Command Feed-forward (CFF). The forward catheter model is implemented for state estimation based on the measured motor outputs. The inverse catheter model developed to allow the controller to operate in state space is as follows. Taking the catheter state x as input with the constant curvature assumption, the coordinates of point C in Figure 3 are derived geometrically as( XC ZC ) =  ( X2+Y 2+Z2 2(X2+Z2) ) X( X2+Y 2+Z2 2(X2+Z2) ) Z  (15) With these coordinate values known the articulation angle \u03c6 is determined from \u03c6 = arctan ( Y\u221a (X \u2212XC)2 + (Z \u2212 ZC)2 ) (16) The geometric configuration parameters are then found as s \u03ba \u03b2  =  \u221a X2+Y 2+Z2 2(1\u2212cos\u03c6) \u03c6\u221a 2(1\u2212cos\u03c6) X2+Y 2+Z2 arctan (\u2212Z X )  (17) With 5 control inputs there is no unique solution for the configuration vector qc without defining additional constraints. In [8], a minimum tension was maintained in the tendon control wires for any given configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000992_s11044-013-9388-1-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000992_s11044-013-9388-1-Figure11-1.png",
        "caption": "Fig. 11 Initial postures of pendulum: qI = (q1, q2) \u2208 {(0,0), (\u03c0,0), (0,\u03c0), (\u03c0,\u03c0)}",
        "texts": [
            "2 Results Figure 9 shows a cross-section of the partitioned phase spaces for configuration parameters (q1, q2) = (\u2212 \u03c0 10 \u223c 0, 9 10\u03c0 \u223c \u03c0). The \u201cnonuniform\u201d partition in Fig. 9 presents the results from the proposed partitioning method. Figure 10 provides the relationships between the number of cells and the least average time. The graphs present the effectiveness of each method, because time is selected as optimality criterion while operating RASMO. The number of cells is described on a logarithmic scale (log4) because the targeted phase space has 4 dimensions. Figure 10a\u2013d present data from different initial states given in Fig. 11. From Fig. 10, it is obvious that the proposed method of partitioning is better than the uniform method because this new approach achieves better motion. Motion quality bias provides useful insights. When the initial state of the pendulum is an inverted form (q1 = 0 in a and c of Fig. 11), the proposed partition is effective (average time was short) at higher resolution. When the initial state of the pendulum is an extended form (q2 = 0 in a and b of Fig. 11), the proposed partition is effective at lower resolution. This may be because effectiveness at higher resolutions is related to the potential energy of the machine (a and c in Fig. 11). In addition, the effectiveness at lower resolutions may be related to effect of the large moment of inertia about the initial state (a and b in Fig. 11). Table 4 summarizes the CPU time of calculating each procedure. The CPU time of calculating the single motion planning from the initial state to the goal state is approximately 200 times less than that of partitioning. From this result, it is expected that a machine (e.g., industrial or space robots) calculates higher quality of motions if once nonuniform partition- Table 4 Calculation cost Resolution 12 16 20 24 28 Partitioning [s] 576 1865 4452 10158 19369 (Subroutine A) [s] 265 857 2017 4666 8284 Motion planning [s] 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002381_eej.22882-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002381_eej.22882-Figure1-1.png",
        "caption": "Fig. 1. Airframe of quadrotor. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]",
        "texts": [
            "1 Mechanism of multicopter In a quadrotor-type UAV, moment produced due to difference between diagonally placed rotors is used for motion in horizontal plane. Vertical motion is determined by total trust generated by the rotors. Counter torques occurring upon rotor rotation cancel each other because the front and rear rotors rotate clockwise, while the left and right rotors rotate counterclockwise. C \u20dd 2016 Wiley Periodicals, Inc. 2.2 Airframe motion equations We consider a quadrotor-type UAV shown in Fig. 1. Here, the centroid is set at B in the body coordinate, and at arbitrary point E in the earth coordinate. Considering moments surrounding axes and forces applied to body coordinate axes, equations of translational and rotational motions of the controlled plant can be expressed as shown below based on Newton\u2013Euler method [5]. \u03d5\u0308 = \u03b8\u0307\u03c8\u0307 Ixx (Iyy \u2212 Izz) \u2212 Jr Ixx \u03b8\u0307 \u03a9r + l Ixx U2, (1) \u03b8\u0308 = \u03d5\u0307\u03c8\u0307 Iyy (Izz \u2212 Ixx ) + Jr Iyy \u03d5\u0307 \u03a9r + l Iyy U3, (2) \u03c8\u0308 = \u03c8\u0307\u03b8\u0307 Izz (Ixx \u2212 Iyy) + 1 Izz U4, (3) x\u0308 = \u2212gsin\u03b8 + (sin\u03d5sin\u03c8 + cos\u03d5sin\u03b8cos\u03c8) 1 m U1, (4) y\u0308 = gcos\u03d5sin\u03b8 + (sin\u03b8sin\u03c8cos\u03d5 \u2212 sin \u03d5cos\u03c8) 1 m U1, (5) z\u0308 = gcos\u03d5 cos \u03b8 + (cos\u03d5cos\u03b8) 1 m U1, (6) where \u03d5, \u03b8, \u03c8 are rotation angles surrounding body coordinate axes, following the order of Euler angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001486_wac.2014.6935970-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001486_wac.2014.6935970-Figure5-1.png",
        "caption": "Fig. 5. Definition of simulator 's variables. (a) Sagittal plane from right side of the robot. (1; is the joint-angle of the robot or a lead screw angle of the crutch. Subscripts Ll, L2, L3, Rl, R2, R3, Lsh and Rsh are the left ankle, left knee, left hip, right ankle, right knee, right hip, left shoulder and right shoulder joints, respectively. Subscripts Le and Rc are the left and right lead screws, respectively. Xi is the position of the foot or the crutch from the hip joint. Subscripts If, r f, Ie and re are the left foot, right foot, left crutch and right crutch, respectively. (b) Frontal plane of the robot. Inclining to the side of the left leg is the positive direction of the roll angle. Yi and Zi are the position of the foot and the crutch from the center of left and right hip joints, respectively. Com is a ground position of the CoM of the robot. The simulator has six DoFs and five links at the lower body, and four OoFs at the upper body including the crutch.",
        "texts": [
            " MOTION SI MULATOR OF PARAPLEGIC PATIENT BO D Y AND WAL K SUPPORT S Y STE M Figure 4 shows the system architecture of the motion simulator of a paraplegic patient body and the walk support system. Specifications of the motion simulator, acceleration sensor and gyro sensor are on Table II, ill and N, respectively. The wa1king robot simulates the patient body wearing an underactuated walk support system on the lower limb. Besides, the robot is operated only by commands from the finger mounted walk controller. Figure 5 shows a definition of system parameters and variables. The robot consists of seven links at the lower body, five links at the upper body. Each of joints, hip, knee, ankle and shoulder has one DoF and rotates around a pitch axis. The upper body simulates a patient who holds crutches in his/her hand. Crutches of the robot have one DoF and are extended/shortened by rotating ball screws with a DC motor. Each joint angle and crutch length is measured by a rotary encoder mounted in a DC motor. The sensor box is mounted on the hip, which measures posture of an upper body by using inclinometers, gyro sensors and the acceleration sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003095_j.cattod.2020.05.059-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003095_j.cattod.2020.05.059-Figure1-1.png",
        "caption": "Fig. 1. The electrochemical cell without gaskets: 1) reinforcement plates, 2) gas distributor, 3) levelling layer, 4) cathode holder, 5) GDE current collector, 6) GDE, 7) catholyte compartment, 8) static mixer, 9) RE connection, 10) CEM, 11) iridium felt on titanium sinter plate, 12) anode current collector with anolyte distributor, 13) anode holder and 14) power connections.",
        "texts": [
            " We discuss the potential of the coupled process experimentally and with modeling results. This study has implications beyond the conversion of cellooligomers to glucose via the electro-Fenton: We present a promising approach for the continuous operation of electro-Fenton processes. All error bars reported in this study represent the standard deviation of experiments which were conducted at least in duplicate. A self-made electrochemical plate and frame reactor was used for the electro-Fenton process. The reactor is sketched in Fig. 1 and consisted of three compartments: 1) a gas supply compartment for oxygen supply to the cathode, 2) a catholyte compartment for the electroFenton reaction and 3) an anolyte compartment for the oxygen evolution reaction. Similar reactors are frequently employed in electrochemical processes with gas diffusion electrodes [31,32]. The catholyte and anolyte were separated by a cation exchange membrane (Fumapem F-14100, Fumatech). A carbon paper was used as the gas diffusion electrode (GDE) on the cathode side (QuinTech Fuel Cell Technology, Freudenberg, H2315 I2C6) and an iridium-coated titanium felt (Bekaert ST/Ti/20/150/85) was used as the anode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003181_tte.2020.3004734-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003181_tte.2020.3004734-Figure1-1.png",
        "caption": "Fig. 1. The topology of the investigated HMC-VFMM [4].",
        "texts": [
            " Besides, a bipolar voltage injection method for the polarity detection of the machine is subsequently expounded. The voltage pulse magnitude determination of the bipolar voltage injection method is discussed. Afterwards, the implementation of the d-axis current pulse injection method based on the estimated rotor position to elevate the MS is introduced. In Section VI, some experiments on a HMC-VFMM prototype are carried out to validate the effectiveness of the proposed MS initialization control scheme. The topology of the investigated HMC-VFMM is shown in Fig. 1. The hybrid parallel/series magnetic circuits are designed as a separated dual-layer PM configuration, in which the parallel circuit can make the majority of the magnetic fluxes produced by NdFeB short-circuited within the rotor core, resulting in a wide flux adjusting range. Additionally, due to the dual-layer structure, the drawbacks of high magnetizing current level and the restricted flux regulation range of the conventional series type [4] are resolved as well. Besides, the series circuit provides a better on-load demagnetization withstand capability, which is beneficial for the online dynamic performance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001794_j.wear.2015.01.053-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001794_j.wear.2015.01.053-Figure3-1.png",
        "caption": "Fig. 3. Assembly drawing (a) and a photograph of the gearbox in use (b).",
        "texts": [
            " Modulus of the ZN type gearing is 4 mm, the centre distance 90 mm, and the pressure angle in the normal plane is 201. Each sixth tooth on the wheel is marked with a corresponding number, which gives a total of 6 teeth on each wheel. The gearbox used in this investigation is selected from a regular assortment after being carefully checked for geometrical accuracy in order to eliminate a potential influence of serious manufacturing errors. The casing of the gearbox is adopted and equipped with temperature sensors, a robust camera carrier (Fig. 3, left, parts 10\u201315) and a precisely built and assembled worm wheel locking mechanism (Fig. 3, left, parts 1\u20137). Both the carrier and the locking mechanism are firmly attached to the casing. The carrier is positioned so that the axis of the camera objective corresponds with the tooth flank normal at the pitch diameter and half face width. Because of the design of additional gearbox equipment and the nature of investigation, the disassembly of the gearbox is not needed. In this way, all the parts maintain their original positions from the very beginning to the end of testing. 2.2. Lubricant Oil used for lubrication of gearing complies with a recommendation for the selection of oils for worm gearboxes based on known data for the rotational speed and nominal power",
            " The quality of this oil complies with ISO 6743-6 (CKC), AGMA 9005 (5EP) and DIN 51517 Part 3 (CLP). Properties of the oil are given in Table 2. In order to prevent abrasive damage due to expected pitting damage fallouts, a closed lubricating circle with oil filtration is used. The applied method consists of two parts: (a) digital image acquisition and (b) digital image processing. In order to maintain same relative position between the camera objective and the observed wheel tooth flank, a specially built locking mechanism is used. The tip of the precisely guided sliding bar (Fig. 3, left, part 4) is pushed between two adjacent teeth of the wheel, always to the same depth. The wheel is then gently loaded to align the worn side of one tooth with the spherical tip of the sliding bar. Assuming an approximately equal amount of sliding wear on all teeth (which is a reasonable assumption for a regular worm pair), the relative position between the observed tooth flank and the camera objective can be assumed equal. Prior to taking a picture, when the position of the wheel is immobilised, the surface of the observed tooth flank is carefully cleaned out from oil. The pitted area is in a sharp contrast to the rest of the surface which is polished due to sliding abrasive motion. In order to prevent a flashing effect on a polished surface and provide uniform, dispersed light on the tooth flank, the top of the gearbox is covered with a piece of tracing paper (Fig. 3, right). A hole in the middle of the tracing paper gives a clear view to the camera objective. A 2576 1926 pixel, 90 dpi colour image is recorded in a lossless raw format [14] which gives the opportunity of software processing before the final image creation. The taken raw image is processed into another lossless image format (tagged image file format, TIFF). The actual tooth flank area is tailored with an image editing tool and processed with a computer vision application. In this most sensitive part of processing, a procedure is created, based on visual data, to automatically distinguish the pitting areas from the rest of the tooth surface, to index and quantify them, and to produce an image with outline boundaries of the tooth as well as the damaged areas"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002402_aim.2016.7576906-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002402_aim.2016.7576906-Figure1-1.png",
        "caption": "Fig. 1. Geometric description of PN guidance",
        "texts": [
            " It is also assumed that the inner-loop control of the interceptor is perfect such that the interceptor can be described by { r\u0307I = vI, v\u0307I = aI , (2) where rI , vI and aI are the position, velocity and acceleration of the interceptor, respectively. The relative position of the target from the interceptor is defined by r , rT \u2212 rI = [x, y, z]T , where superscript \u00b7T denotes the transpose of vector. The relative velocity with respect to the target and the interceptor can be defined by v , vT \u2212 vI, and it follows that { r\u0307 = v, v\u0307 = aT \u2212 aI. (3) As is displayed by Fig. 1, the principle of PN is to maintain the direction of the relative position and reduce the relative distance. Geometrically, it can be expressed by vT sin\u03b8 \u2212 vI sin\u03b4 = 0, (4) which is to keep the direction of relative position, and r\u0307 = vT cos\u03b8 \u2212 vI cos\u03b4 < 0, (5) where r , \u2016r\u2016, indicating that the relative distance decreases. Assumption 1: The inner (or attitude) loop control of the interceptor rotary UAV is designed and implemented perfectly, such that the acceleration of the interceptor can be regarded as the guidance input"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002085_j.wear.2016.04.027-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002085_j.wear.2016.04.027-Figure1-1.png",
        "caption": "Fig. 1. Small hole with radius R in a large plate subjected to an alternating nominal stress, S.",
        "texts": [
            " Note however that also these approaches typically require either full elastoplastic stress Please cite this article as: A. Ekberg, et al., Stress gradient effects in s (2016), http://dx.doi.org/10.1016/j.wear.2016.04.027i analyses or the identification of a stress concentration factor, which is not straightforward in RCF applications. To establish a case of plain fatigue that can be contrasted to the case of rolling contact fatigue, we study stresses in the vicinity of a small round hole in the centre of a large, thin plate. This stress field can (see e.g. [6]) with nomenclature following Fig. 1 be expressed in polar coordinates as \u03c3r \u00bc S 2 1 R r 2 ! \u00feS 2 1 4 R r 2 \u00fe3 R r 4 ! cos 2\u03b8 \u03c3\u03b8 \u00bc S 2 1\u00fe R r 2 ! S 2 1\u00fe3 R r 4 ! cos 2\u03b8 \u03c4r\u03b8 \u00bc S 2 1 3R4 r4 \u00fe2R2 r2 ! sin 2\u03b8 \u00f03\u00de To quantify this multiaxial state of stress in the vicinity of the hole, the Dang Van equivalent stress [7] is employed. The equivalent stress is defined as \u03c3dv \u00bcmax t \u03c3d 1;a t\u00f0 \u00de \u03c3d 3;a t\u00f0 \u00de 2 \u00fecdv\u03c3h t\u00f0 \u00de ! \u00f04\u00de Here \u03c3d 1;a and \u03c3d 1;a are deviations from the mid value over a stress cycle of the largest and smallest principal component of the deviatoric stress tensor, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000611_978-3-319-48194-4_37-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000611_978-3-319-48194-4_37-Figure1-1.png",
        "caption": "Figure 1: Schematic diagram of a laser cladding setup and model-part domains",
        "texts": [
            "13 Numerical Modeling The numerical model presented in this work is based on computational fluid dynamics methods (CFD) for the powder flow rate model and the melt pool dynamics model and on finite element method (FEM) for the solidification and metallurgical models. The different methods are coupled into an integrated model to predict the melt pool temperature and dimensions, cooling rates and resulting metallurgical phases after solidification of the liquid metal from initial process parameters. A schematic diagram of the modeling domains is shown in Figure 1. with laser irradiance and the weld pool physics was summarized in [18, 23]. The coupling of CFD and FEM models was described in [22, 24]. Due to space limitations, this paper focusses on the prediction of the metallurgical properties and the resultant residual stresses. The metallurgical model inside the FEM system is based on the model of Leblond [20]: ( ) ( \u0307) ( ) ( ) (1) with P phase proportion, T temperature, t time, heating or cooling rate, Peq phase proportion at equilibrium and \ud835\udf0f a delay time as a function of temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002615_icsens.2016.7808663-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002615_icsens.2016.7808663-Figure1-1.png",
        "caption": "Fig. 1. Schematic picture of gold interdigitated electrodes with configuration of 50 pairs of fingers, thickness t = 120nm, width w = 10 m, distance between each digit g = 10 m and length line of l = 3 mm.",
        "texts": [
            " EXPERIMENTAL Poly(ethyleneimine) (PEI), poly(styrenesulfonate) (PSS), poly(3,4-ethylenedioxythiophene):poly(styrenesulfonate) (PEDOT:PSS), poly(pyrrole) (PPy), poly(aniline) (PANI), chitosan extracted from shrimp (Chit), 3,4- ethylenedioxythiophene monomer (EDOT), copper (II) (CuTsPc) and nickel (II) (NiTsPc) tetrasulphonated phthalocyanines were purchased from Sigma-Aldrich. Solutions of KCl, HCl and glucose from Aldrich were prepared using ultrapure water from a Millipore Direct-Q5 system with concentrations varying from 10-7 to 10-4 mol/L. Fig. 1 shows the schematic picture of gold interdigitated electrodes (IDEs) fabricated at LNNano/LNLS laboratory using conventional photolithographic technique. Surface modification with nanostructured films on gold interdigitated electrodes was carried out using the layer-by- 978-1-4799-8287-5/16/$31.00 \u00a92016 IEEE layer (LbL) and electrochemical deposition techniques. All solutions for adsorption of the films were prepared using ultrapure water from a Millipore Direct-Q5 system with concentration of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001272_imece2013-64696-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001272_imece2013-64696-Figure2-1.png",
        "caption": "FIGURE 2. SCHEMATIC OF THE SPECIMEN WHERE b DENOTES THE WIDTH, l THE LENGTH AND h THE HEIGHT OF THE SPECIMEN.",
        "texts": [],
        "surrounding_texts": [
            "To avoid the high precision requirements for the production of small kinematic structures such as laparoscopic grippers, compliant mechanisms and flexible hinges are employed [2]. Even whole medical robots [5] can be designed completely out of flexible elements. For those applications the flexural modulus is essential, because the deflection of the hinges and elements needs to be known. Experiments already published concerning the calculation of thin flexural structures could not verify the predictions calculated with the known flexural modulus [2]. They suggested, that there is a dependency of the modulus on the thickness of the specimen. This paper researches this dependency."
        ]
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure13-1.png",
        "caption": "Fig. 13: Incremental TH simulation field solution for the 5th stator belt harmonic. The same procedure is repeated till the 13th.",
        "texts": [
            " 2) TH Simulations: in this stage the rotor reaction to the stator belt harmonics has to be determined, performing frequency domain simulations. In the TH problem, the field harmonic is reproduced properly imposing the current in the stator slots; in particular, the real part of the current density distribution has to produce a magnetic field in the same direction that the stator harmonic has in the considered time instant in the MS simulation. The imaginary part is imposed just to consider the rotation of the field with respect the rotor. The field solution is shown in Fig. 13 and it is possible to store the induced current in each portion of each rotor slot, to impose them later on the MS simulation. Authorized licensed use limited to: Carleton University. Downloaded on June 30,2020 at 13:07:46 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. 3) Final MS Simulation: the complete field solution is obtained supplying the harmonic windings according to the previous TH simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000775_s1052618814010105-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000775_s1052618814010105-Figure1-1.png",
        "caption": "Fig. 1.",
        "texts": [
            " The second equation is the equation of small elastic vibrations {\u0394qd}, which appear in the hinges of manipulation systems. The right side of this equation reflects elastic and dissipative properties of hinges. The third equation, according to (5), associates the quasi statically small elastic deformations {\u0394qks} and forces developed by drives. Let us compare the results of numerical modeling of the dynamics of manipulation systems with elastic hinges found using mathematical model (4), (5), and (8). We will perform modeling by the example of the three link manipulation system (Fig. 1). The first link of the manipulation system under study is modeled by a thin wall pipe with length l1, radius R1, and mass m1, which is rotated around the vertical axis coin ciding with the central longitudinal axis of the pipe. The second link is pivotally connected with the first M[ ] q \u0394q+( ) M[ ] q( ) \u2202 M[ ] \u2202ql \u0394ql, S[ ] q \u0394q+( ) l 1= n \u2211+ S[ ] q( ) \u2202 S[ ] \u2202ql \u0394ql, l 1= n \u2211+= = K[ ] q \u0394q+( ) K[ ] q( ) \u2202 K[ ] \u2202ql \u0394ql. l 1= n \u2211+= M q \u0394q+( )[ ] M q( )[ ] O \u0394q( ), S q \u0394q+( )[ ]+ S q( )[ ] O \u0394q( ),+= = K q \u0394q+( )[ ] K q( )[ ] O \u0394q( )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002778_978-3-030-29131-0-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002778_978-3-030-29131-0-Figure11-1.png",
        "caption": "Fig. 11 Demonstration of the ACC on a highway scenario",
        "texts": [
            " a Recognition of walking activities and b Recognition of gait phases and events Wearable sensor for sit-to-stand activity\u2014The recognition of sit-to-stand (SiSt) and stand-to-sit (StSi) activities, has also been studied in laboratory environments usingwearable sensors. In the experiment shown in Fig. 10, the participant was asked to perform multiple repetitions of SiSt activity to collect acceleration data from one IMU. These data were processed by a probabilistic approach for recognition of sit, transition and stand states during the SiSt activity [47]. In addition,multiple segments that compose the transition state were recognised, which is important to achieve a robust and accurate control of assistive robots (Fig. 11). Other aspects that need to be considered in a real or outdoor environment could affect the recognition accuracy Fig. 10 Inertial measurement units employed for recognition of sit-to-stand and stand-to-sit activities Assistive Gait Wearable Robots \u2026 89 of SiSt. Some of these aspects are the delays in the pre-processing steps to smooth the signal, type of chair, optimal location of sensors and speed to move from sit to stand. Challengeswithwearablemeasurement systems\u2014Wearable sensors have been successfully used for recognition of human activities",
            " For this particular test case, a section of E40 highway scenario near Bertem in Belgium is virtualised using real world map data. The virtual scenario, sensor models and traffic simulation are all developed using Simcenter\u00ae Prescan. A dedicated button on the steering wheel is used for enabling/disabling the ACC functionality. In this scenario, the ego vehicle is cruising at approximately 100 km/h and approaching a leading vehicle. To avoid the collision, the ego vehicle slows down to the same speed as the leading vehicle as is shown in Fig. 11. After 41 s, the ego vehicle decides to overtake the leading vehicle. From this moment, the ACC increases the velocity of the ego to the previously defined setpoint. Autonomous Intersection Crossing The autonomous intersection crossing controller is demonstrated in the second scenario. The developed control algorithm is validated by having the ego vehicle approaching a cross-road together with one other road user. At a certain moment, their trajectories are overlapping and a collision might occur",
            " 10 Simulated contour error at the encoders and the TCP The simulation time for this contour is approximately 30 s. That is very efficient, although not yet real-time. In manufacturing technology, process stability and productivity tend to be the most vital properties for success. Coupling of a machine tool covering the whole mechatronic system with a production process model represents the top level of analysis. This is only possible with efficient, accurate, and low-order models. Kuffa [21] showed, by means of a MORe model of a test rig for high performance dry grinding as shown in Fig. 11, that this coupling is feasible. He divided the surface profile calculation in a machine-dependent part and a kinematic roughness part, which depends on the geometry of the grinding grains. Themachine dependent behaviourwas analysed bymeans of a simplified grinding process model. Good correspondence of the surface profile between measurements and simulation was achieved as shown in Fig. 12. An importantmeasure in precisionmachining is the volumetric accuracy of amachine tool. Volumetric accuracy is, according to Ibaraki and Knapp [22], represented by a map of position and orientation error vectors of the tool over the volume of interest"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002223_978-3-319-22894-5_19-Figure19.4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002223_978-3-319-22894-5_19-Figure19.4-1.png",
        "caption": "Fig. 19.4 Polyacrylate-based side-chain LCE actuators. (a) Alternated UV-visible photo-isomerization at different points of the azobenzene-containing LCE actuator ring which creates a local contraction and a local expansion of the belt and rotates the pulleys; (b) Micropump based on the bending motion of a LCE film which contains azobenzene moieties which upon UV-irradiation bends and with visible light relaxes back. (c) Sequence of thermo-resistors actuation showing changes in length and transparency of the LCE material upon heating and cooling under the application of 320 mW of electrical power",
        "texts": [
            " The main actuation principle relies on the trans-to-cis disorder-induced photo-isomerization upon UV-irradiation, and the order-induced cis-to-trans photo-back-isomerization with visible light. By the simultaneous irradiation using both UV and visible light in opposite sides of a LCE ring, the combination of intermittent contraction and expansion movements results in a rolling motion of the film, which drives the two pulleys forming the motor. A scheme showing the working principle of the device is depicted in Fig. 19.4a, where the main elements forming the device can be identified. An example of the photoinduced motion of the motor in a counterclockwise direction is depicted in the sequence of images below. In order to reinforce the LCE film and thus improve its mechanical properties, a 50 \u03bcm thick flexible polyethylene (PE) sheet was attached on the photo-crosslinked azobenzene-containing LCE ribbon by thermal bonding creating a 68 \u03bcm thick laminated film, which was used to form the motor belt by connecting both ends",
            " Similar laminated films were tested upon UV exposure at different intensities showing generation of mechanical force by photo-irradiation. Afterward, Chen et al. (2010) developed a light-driven micropump by incorporating photo-isomerizable azobenzene moieties into the polyacrylate-based sidechain LC material. The reversible bending behavior of this material upon UV and visible light exposure is the responsible for inducing the membrane\u2019s movement and to pump a fluid. The working principle of the micropump, as well as a picture of the experimental prototype, is shown in Fig. 19.4c. Upon irradiation with UV light, the contraction gradient through the thickness of the film induces a downward bending which results in the reduction of the pump chamber volume, and the corresponding generation of pressure. In this way, the fluid in the chamber is forced to go to the pipe outlet. When the sample is exposed to visible light, a recovery of the film flatness is achieved by the upward bending of the film, and reduces the chamber\u2019s pressure which stops the flow. Similar to the previously described microgripper (Sa\u0301nchez-Ferrer et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000061_apusncursinrsm.2019.8889294-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000061_apusncursinrsm.2019.8889294-Figure1-1.png",
        "caption": "Fig. 1. Miura unit cell.",
        "texts": [
            " The results indicate a gain variation from 4 dB to 8 dB in the frequency range of 686 MHz to 760 MHz. A Yagi-Uda dipole antenna with one reflecting element and four directing elements is designed to operate in the 698MHz to 806 MHz band, [5]. Deployability and reconfigurablity are achieved by mounting the antenna on a periodic origami This work was supported by the Air Force Office of Scientific Research under grant FA9550-18-1-0191, the National Science Foundation under Grant EFRI 1332348, and the FIU Presidential fellowship. structure. Fig. 1 shows the unit cell used for this design which was initally proposed by Miura et al. [6]. The unit cell has a volume of (2L + V ) \u00d7 2S \u00d7 H where L, V , S, and H are given in terms of the edge lengths a and b, the fold angle \u03b8, and the vertex angle \u03b3 as defined in (1) to (4): H = a \u00b7 sin(\u03b8) \u00b7 sin(\u03b3) (1) S = b \u00b7 cos(\u03b8)tan(\u03b3)\u221a 1 + cos(\u03b8)2tan(\u03b3)2 (2) L = a \u221a 1\u2212 sin(\u03b8)2sin(\u03b3)2 (3) V = b\u221a 1 + cos(\u03b8)2tan(\u03b3)2 (4) To adjust the dimensions of the unit cell to the antenna size, the angles \u03b3 and \u03b8 are fixed at 45o. A Uniform spacing of the elements along the x-axis of the unit cell, as shown in Fig. 1, allows b to be found since the inter-element spacing is equal to 3S. By orienting the elements\u2019 axis parallel to the y-axis of the unit cell, the length a can be found in terms of the largest element length. For the case of the Yagi-Uda, the reflector is the largest element and is set here equal to 4L. A rigidly-foldable cylindrical origami structure is created from the previously designed unit cell using the method 453978-1-7281-0692-2/19/$31.00 \u00a92019 IEEE AP-S 2019 outlined in [4]. The antenna is mounted to this structure as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003821_icece51571.2020.9393056-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003821_icece51571.2020.9393056-Figure5-1.png",
        "caption": "Fig. 5 Increment of Stator flux in an induction motor with the change in stator voltage vector.",
        "texts": [],
        "surrounding_texts": [
            "vector. This reference voltage space vector is synthesized by three stationary voltage vectors. This scheme shows lower current, low torque ripple and good DC-link voltage utilization when compared to other PWM techniques and operates with a constant switching frequency to reduce the switching losses [11]-[13].\n\u00ce *s' 2 5* 2 v sa + v sR )sR\n6 = tan -1\ns* v\ns*\n\\\nsR\nsa\n(1)\n(2)\n*svs\nv\nIn any sector the reference voltage is obtained by selecting two active vectors and one null vector. The time durations of these three vectors should be on in sector1 is given by T1, T2, and T0. T1 is the active time for voltage vector V1. T2 is the active time for voltage vector V2. And TO is the time for Null vector is on.\n73 x Ts x Vs n \\\nT1 =\n- 6 V3 )\nV (3)\nT2 = 43 x Ts x Vs* x sin (6)\nV (4)\nT0 = Ts - T1 - T2 (5)\nIn any sector the general duty times for these three vectors is given by equations (5), (6) and (7).\n73 x Ts x V f * x sin\nT1 =\n( n n 3 - - 6 + (K - l ) x - V 3 3 ) V\u201e (6)\n73 x Ts x Vss* x sin 6 - ( K - 1):\nT2 =\nn\n3 )\nV, (7)\nFor linear range of operation of three phase inverter, the modulation index should lie between zero and one. The reference voltage takes minimum value at modulation index is zero and maximum value at modulation index is one.\nHence the range of modulating reference signal in space vector modulation is given by equation (9).\n3 x Vss*\nV 0 < ma < 1 (8)\nd The maximum line to line RMS voltage value in space vector modulation is given by equation (13). Fig. 3 shows the operation of the sectors of the proposed DTC scheme.\nV,\n3 ^ (9)0 < V s* < ds\nrs*v =43 xs maxi/- l )(rms )\n)\n2 ,s*Vs = 0.707 x V,s max(l- l ) d(rms)\n(11)\n(12)\nFig.3 Sectors of operation in DTC-SVM with reference space vector vref In the closed loop torque control the actual value of torque is compared with reference torque and the error produced is given as input to torque PI controller whose output is considered as q-axis of stator reference voltage [14]. Similarly closed loop flux control is operated and the output is considered as d-axis of stator reference voltage. The control voltage signals in d, q axis is converted to a, ft axis. From a, ft reference voltages, the control signal is calculated which is considered as reference voltage space vector. This reference voltage space vector controls the stator flux of an induction motor and hence the torque [15]. Fig. 4 shows closed loop torque and Flux control with suitable gain values.\n404\nAuthorized licensed use limited to: National University of Singapore. Downloaded on July 03,2021 at 22:37:30 UTC from IEEE Xplore. Restrictions apply.",
            "360\u00b0 * f AO = --- \u2014 f = 3.6 deg (13)\nf s\nIII. Sim u l a t i o n Re s u l t s Us in g DTC-SVM Te c h n iq u e\nThe proposed DTC-SVM scheme is simulated using MATLAB Simulink. Fig. 6 shows the sector selection from 6 to 1 ( 6 active vectors ) with time. Any one zero vector is taken into account for calculating the total time for the space vector reference as per the space vector philosophy.\nIn order to show clear comparison of current, torque ripples both the schemes are simulated under the same conditions. Motor Specifications are listed are listed in Table I.\n(i) Steady state torque responses are shown in Fig. 7. Fig. 7(a) shows the results of steady state response in a DTCSVM scheme at Full load condition i.e. 80 Nm and rated speed of 1430 rpm. At constant switching frequency of 5 kHz it shows faster steady state response and it has less torque ripple of 3.5 Nm as compared to 5.4 Nm in conventional DTC scheme at the same operating conditions. Fig. 7(b) shows the results of steady state response in a DTCSVM scheme at half load condition i.e. 40 Nm and rated speed of 1430 rpm. It shows faster steady state response and it has less torque ripple of 2.4 Nm. Fig. 7(c) shows the results of steady state response in a DTC-SVM scheme at A th of load condition. It has very less torque ripple of 2 Nm. This work proved that there is an overall reduction in the torque ripple when DTC-SVM is used for control of IM drive. (ii) Transient state Torque responses in DTC-SVM scheme is shown in Fig. 8. Fig. 8(a) shows the transient state response of a DTC-SVM scheme at Full load condition. It settles in 30 ms. The controller is said to be robust. The conventional DTC takes 150 ms which is very high. Hence this controller can replace the conventional one for high torque loads. Similarly for half load and one-fourth load conditions the transient responses are found to be good with very less settling time of 18 ms and 15 ms as shown in Fig. 8(b), 8(c)\nrespectively. The switching pulses for the IGBT switches of the inverter circuit operating at 5 kHz is shown in Fig. 9. At transient state condition, phase current waveforms of Three phase induction motor in DTC-SVM scheme are shown in Fig. 10. Fig. 10 shows steady state current in DTC-SVM scheme at half load condition i.e. 40 Nm and rated speed of 1430 rpm. (i)\n405\nAuthorized licensed use limited to: National University of Singapore. Downloaded on July 03,2021 at 22:37:30 UTC from IEEE Xplore. Restrictions apply.",
            "operation. The study shows that the control scheme is more advantageous in terms o f reduction in current and torque ripples by nearly 40% and 35% in an induction motor.\nHence this scheme o f control can be used fo r the IM drive in\nE V applications. The TH D in the stator current is less as per the standards. The dynamic nature o f the D TC control\nscheme is achieved fo r rated torque reversal o f 80 Nm. The drawback o f conventional D TC is elim inated by using DTC - S V M scheme w ith low er current and torque ripples during\nboth steady state and transient conditions. I t has a faster response as w e ll w ith good dynamic characteristics to command values o f Torque and flux. Better u tiliza tion o f the\ndc lin k voltage is obtained using D TC w ith SVM . The future w ork includes w orking w ith less number o f PI controllers to achieve IM good performance.\nRe f e r e n c e s\n[1] D. Casadei, F. Profumo, G. Serra and A. Tani, \u201cFOC and DTC: two viable schemes for induction motors torque control,\u201d in IEEE Transactions on Power Electronics, vol. 17, no. 5, pp. 779-787, Sept. 2002, doi: 10.1109/TPEL.2002.802183. [2] T. Noguchi and I. Takahashi, \u201cQuick torque response control of an induction motor based on a new concept,\u201d IEEJ Tech. Meet. on Rotating Machine RM84-76, pp. 61-70, Sep. 1984. [3] M. Hafeez and M. NasirUddin, \u201cA new torque hysteresis control algorithm for direct torque control of an iM drive,\u201d 2011 IEEE International Electric Machines & Drives Conference (IEMDC), Niagara Falls, 2011, pp. 759-764. [4] M. Masmoudi, B. El Badsi and A. Masmoudi, \u201cDirect Torque Control of Brushless DC Motor Drives With Improved Reliability,\u201d in IEEE Transactions on Industry Applications, vol. 50, no. 6, pp. 3744-3753, Nov.-Dec. 2014, doi: 10.1109/TIA.2014.2313700. [5] A. Jidin, N. R. N. Idris, A. H. M. Yatim, T. Sutikno and M. E. Elbuluk, \u201cAn Optimized Switching Strategy for Quick Dynamic Torque Control in DTC-Hysteresis-Based Induction Machines,\u201d in IEEE Transactions on Industrial Electronics, vol. 58, no. 8, pp. 3391-3400, Aug. 2011, doi: 10.1109/TIE.2010.2087299. [6] C. Lascu, I. Boldea and F. Blaabjerg, \u201cA modified direct torque control (DTC) for induction motor sensorless drive,\u201d Conference Record o f 1998 IEEE Industry Applications Conference. Thirty-Third IAS Annual Meeting (Cat.No.98CH36242), St. Louis, MO, USA, 1998, pp.415-422. [7] T. G. Habetler, F. Profumo and M. Pastorelli, \u201cDirect torque control of induction machines over a wide speed range,\u201d Conference Record of the 1992 IEEE Industry Applications Society Annual Meeting, Houston, TX, USA, 1992, pp. 600-606. [8] T. G. Habetler, F. Profumo, M. Pastorelli and L. M. Tolbert, \"Direct torque control of induction machines using space vector modulation,\" in IEEE Transactions on Industry Applications, vol. 28, no. 5, pp. 1045-1053, Sept.-Oct. 1992, doi: 10.1109/28.158828. [9] D. Casadei, G. Grandi, G. Serra and A. Tani, \u201cEffects of flux and torque hysteresis band amplitude in direct torque control of induction machines,\u201d Proceedings o f IECON'94 - 20th Annual Conference of IEEE Industrial Electronics, Bologna, Italy, 1994, pp. 299-304... [10] E. Ozkop and H. I. Okumus, \u201cDirect torque control of induction motor using space vector modulation (SVM-DTC),\u201d 2008 12th International Middle-East Power System Conference, Aswan, 2008, pp. 368-372, doi: 10.1109/MEPCON.2008.4562350. [11] D. Wang et al., \u201cReduction of Torque and Flux Ripples for Robot Motion Control System Based on Sv M-DTC,\u201d 2018 37th Chinese Control Conference (CCC), Wuhan, 2018, pp. 5572-5576. [12] N. R. N. Idris and A. H. M. Yatim, \u201cDirect torque control of induction machines with constant switching frequency and reduced torque ripple,\u201d in IEEE Transactions on Industrial Electronics, vol. 51, no. 4, pp. 758-767, Aug. 2004. [13] A. K. Singh, A. Dalal and P. Kumar, \u201cAnalysis of induction motor for electric vehicle application based on drive cycle analysis,\u201d 2014 IEEE International Conference on Power Electronics, Drives and Energy Systems (PEDES), Mumbai, 2014, pp. 1-6. [14] Y. Tang and G. Lin, \u201cDirect torque control of induction motor based on self-adaptive PI controller,\u201d 2010 5th International Conference on Computer Science & Education, Hefei, 2010, pp. 1230-1234. [15] H. Moghbeli, M. Zarei and S. S. Mirhoseini, \u201cTransient and steady states analysis of traction motor drive with regenerative braking and using modified direct torque control (SVM-DTC),\u201d The 6th Power Electronics, Drive Systems & Technologies Conference (PEDSTC2015), Tehran, 2015, pp. 615-620.\n406\nAuthorized licensed use limited to: National University of Singapore. Downloaded on July 03,2021 at 22:37:30 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_30_0002408_cdc.2016.7799264-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002408_cdc.2016.7799264-Figure1-1.png",
        "caption": "Fig. 1. 3-DOF Quanser helicopter [6]",
        "texts": [
            " In Section III a state-feedback second order sliding mode controller is proposed, while in Section IV the dynamic output-feedback second order sliding mode controller is discussed, which relies on the exploitation of a suitably designed nonlinear Thau observer for the state of the 3-DOF helicopter. In Section V the proposed residualbased fault detection and isolation scheme is detailed, which includes second-order differentiators exploited to provide the exact time derivatives of the outputs of the system and to build the necessary residual signals. Finally, simulation results are presented in Section VI. Conclusions are drawn in Section VII. The considered system (see Figure 1) is the 3-DOF Quanser helicopter [6], which is widely proposed in literature for experimental setup of control approaches. This system represent an underactuated 3-DOF mechanical system, driven 978-1-5090-1837-6/16/$31.00 \u00a92016 IEEE 6464 by two DC motors. All the parts are connected with revolute joints and the following three variables are analyzed: the elevation angle \u03b5, the pitch angle p and the travel angle \u03bb. See Figure 2 for the considered parameters and variables. This setup consists of a base on which a long arm is mounted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003914_gt2015-43940-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003914_gt2015-43940-Figure2-1.png",
        "caption": "Figure 2. Image showing a) build orientation of DMLS coupons with support structures in red and b) orientation of channel surfaces relative to build direction.",
        "texts": [
            " Each of these five coupons had a row of uniformly sized rectangular channels that varied in width, height, and number of channels. See Table 1 for dimensions. The remaining five coupons (L-1x-In, L-2x-In, M-1x-In, M2x-In, S-2x-In) were made using the same specifications as the five made from CoCr; however, they were made using in-house facilities with a nickel alloy powder that has the same chemical composition as Inconel\u2122 718. Details on the manufacturing of the coupons are given by Snyder, et al. [14]. All DMLS coupons in this study were built at a 45\u00b0 angle from the horizontal as illustrated in Figure 2a. Support structures used for the coupons fabricated in-house are illustrated in the figure as well. The machine operating parameters used for the laser sintering process were calculated using recommendations from the manufacturer for the specific materials used [14,15]. Through the laser sintering process, distortion is a common phenomenon because parts are being subjected to very high thermal gradients during the building process. Warping effects are difficult to predict beforehand and are often determined empirically after a series of test builds",
            " The roughness parameters of each surface for each channel were determined from the point cloud of data using a computer code that was developed in-house. The code calculated the difference between each data point of the CT scan and the fitted surface. Figure 6 shows a plot of the roughness of two surfaces of one channel of the M-2x-In coupon. The image in Figure 6a is a portion of the surface facing upward at a 45\u00b0 angle during the build (i.e. the normal vector of the surface was pointing upward and angled 45\u00b0 from the build direction vector noted in Figure 2b). This surface is noticeably smoother than the surface 5 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use facing downward at a 45\u00b0 angle during the build (Figure 6b). Variation in surface roughness has been reported by many to be a strong function of the build orientation [2\u20134,6]. Generally, the upward-facing surfaces have smaller roughness than downwardfacing surfaces. The height differences from the surface fit were used to calculate an arithmetic roughness average, Ra, which is expressed mathematically in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure3-1.png",
        "caption": "Fig. 3. Comparison of contact surfaces (a) and equivalent cumulated strains (b) with facetized and analytical tool descriptions.",
        "texts": [
            " Given the relative coarseness of the mesh (see Fig. 1b), the contact area is limited to very few nodes, so contact is not detected at the beginning of the process and temporal oscillations are observed. This is likely regarded to come from the coarse and non-smooth description of the cylindrical tools discretized by linear C0 triangular facets. In order to check this hypothesis, an analytical description of the tools is introduced. It is easily applicable to such a process where obstacles are cylinders and cones. Simulation results (see Fig. 3) show that the quality is thus significantly ameliorated. Even though the tools were quite finely discretized (see Fig. 4), the analytical formulation notably improves contact treatment by establishing contact earlier and over a larger area (see left part of Fig. 3). It results into lower, more homogenous and more physical values of the cumulated strains (see right part of Fig. 3). This introductory work shows the importance of smooth description of obstacles for problems with reduced contact area. For more complex shapes, such as encountered in general contact mechanics, the analytical description is not possible. It is then necessary to develop a simple, robust and efficient smoothing method that is compatible with large 3D calculations in a parallel computational environment. In this work frame, contact obstacles are only defined by their finite element surface discretization using linear facets, so providing a C0 continuous interpolation with discontinuous normals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure1.7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure1.7-1.png",
        "caption": "Fig. 1.7 An axial flux permanent magnet motor [21]",
        "texts": [
            "2 Linear machine for electromagnetic catapult [8] . . . . . . . . . . 3 Figure 1.3 Linear machines for applications of logistics [11]. . . . . . . . . 4 Figure 1.4 Linear machines for applications of Industrial Automation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 Figure 1.5 Linear machines for applications of rail transportation . . . . . 6 Figure 1.6 Winding arrangements for tubular PM linear machines [19] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 Figure 1.7 An axial flux permanent magnet motor [21] . . . . . . . . . . . . 9 Figure 1.8 Structure of the moving-coil-type linear DC motor [22] . . . . 10 Figure 1.9 Structure of the moving-coil-type linear DC motor [23] . . . . 10 Figure 1.10 Schematic of reciprocating linear generator system [24] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.11 A linear vibration-driven electromagnetic micro-power generator [26] . . . . . . . . . . . . . . . . . . . . . . . ",
            " Analytical formulas are derived to predict the open-circuit electromotive force, the thrust force, the iron loss, and the winding resistance and inductances, as well as the converter losses. The same magnetization pattern can also 8 1 Introduction be implemented to machine design with different winding phases in Figs. 1.6b, c to reduce EMF harmonics and force ripple. They all have a integer ratio of slot number to pole number, leading to larger cogging force. A double-sided slotted torus axial flux permanent magnet motor suitable for direct drive of electric vehicle is proposed in [21]. Its construction is illustrated in Fig. 1.7. The magnet poles are arranged alternatively along the rotor circumference, and magnetized in axial direction. The motor can be easily mounted compactly onto a vehicle wheel, fitting the wheel rim perfectly. 1.4 Magnet Patterns of Linear Machines 9 Nirei et al. presents the evaluation of a moving-coil-type cylindrical linear DC motor [22]. The structure is illustrated in Fig. 1.8. It consists of a coil, permanent magnets, yokes, a coil bobbin held by the arms of a coil holder and linear bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002390_gt2016-57410-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002390_gt2016-57410-Figure16-1.png",
        "caption": "Figure 16. Retrofit rear bearing with optimized ISFD",
        "texts": [
            " In this particular design, the ISFD bearing was integrated with the same internal design as the original tilting pad bearing, albeit using a 4 pad configuration, instead of 5 pads, to simplify the mechanical design. It must be noted that the internal bearing configuration is not as important to the rotordynamic behavior since the support stiffness is mainly determined by the ISFD stiffness, which is significantly lower than the bearing stiffness. The operation of the subject steam turbine was transformed with the original bearing at the steam admission side and retrofit ISFD bearing at the exhaust end as shown in Figure 16. The \u201cS\u201d shape spring was carefully designed using an internal design code (XLISFD\u2122) to have the selected stiffness value of 4.38E+07 N/m (250 klb/in). In order to have the selected damping value of 2.1E+05 Ns/m (1200 lb-s/in), the damper film gap was designed to be 0.457 mm (0.018 inch) which is much larger compared to conventional SFD. The oil inlet orifices of 2.18 mm (0.086 inch) are located at each 6 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000345_978-94-007-7194-9_2-1-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000345_978-94-007-7194-9_2-1-Figure6-1.png",
        "caption": "Fig. 6 The self-motion of the arm is shown as a rotation of the arm plane, determined by the upper-/lower-arm links, around the line connecting the shoulder and wrist joints. The rotation angle \u02db can be associated with parameter bv in (21)",
        "texts": [
            " Note, however, that in contrast to the velocity-level solution, a zero input for any of these quantities would not terminate the self-motion. This problem can be alleviated by a careful choice of R a or ba, such that the respective inverse kinematic solutions are integrable. Alternatively, a joint damping term can be added for damping out the conserved self-motion velocity. In contrast to a nonredundant limb, a kinematically redundant limb can move even when its end-link is immobilized (V D 0). Such motion is shown in Fig. 6 for the arm; the hand remains fixed w.r.t. the arm root frame, while the elbow rotates around the line connecting the shoulder and wrist joints. Such type of motion is known as self-motion, internal motion, or null motion. Self-motion is generated with the joint velocity obtained from the following homogeneous differential relation: J . / P D 0; P \u00a4 0: (25) Since n > 6, the Jacobian is non-square (6 n), and the above equation is characterized as an underdetermined linear system. Hence, there is an infinite set of solutions, each nontrivial solution representing a self-motion joint velocity: f P n W P D N ",
            " On the other hand, there is a class of tasks wherein the posture of the humanoid is allowed to variate, i.e., a whole-body motion is admissible and also desirable. One such task is arm reach, quite often used as a benchmark task (see, e.g., [41, 54\u201356]). Thereby, the whole-body posture variation is employed to enlarge the workspace of the arm. With this task, as well as with other similar motion tasks, the motion of the hand(s) is specified w.r.t. the inertial frame fW g. This means that motions in the joints along the entire kinematic chain will contribute to the hand motion. In the example in Fig. 6, the total number of joints becomes ntotal D nleg C nspine C narm D 6C 1C 7 D 17. Thus, the degree of redundancy is 17 6 D 11. With such a high degree of redundancy, it is possible to realize multiple \u201cadditional\u201d tasks. Indeed, besides the hand motion tasks, the robot has to simultaneously perform a number of other motion tasks, e.g., to keep balance while avoiding obstacles, selfcollisions, singularities, and joint limits and to follow a moving object visually: the so-called gaze task. All these tasks impose constraints on the motion, referred to as motion-task constraints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure5-1.png",
        "caption": "Fig. 5. Proposed 12S-13P E-Core CPSFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " Analysis reveals that conventional 6S-10P E-Core SFPMM and 6S-10P C-Core SFPMM retain same PM volume as that of 12S-10P E-Core SFPMM and increased slot area, however author fails to compensate effects of leakage flux. Until now, many researchers tried to reduced PM as much as possible and suppress flux leakage but unfortunately both effects are not considered at a time. In this paper, alternate Consequent Pole SFPMM (CPSFPMM) with partitioned PMs are introduced which further reduce PM usage and suppress flux leakage completely. For analysis and electromagnetic performance evaluation, this paper proposed three topologies of CPSFPMM (as shown in Fig. 5, Fig. 6, Fig. 7) corresponding to three topologies of conventional SFPMM as listed in table I. Based on Finite Element Analysis (FEA) proposed model has successfully reduced PM usage much more and suppress PM leakages from the end and enhance flux modulation effect. Hence proposed model reduces machine cost furthermore, retaining electromagnetic performance. The rest of the paper is organized as, section II present design parameter, design methodology and working principle of proposed CPSFPMM, section III illustrates FEA based electromagnetic performance analysis and finally, some conclusions are drawn in section IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000548_978-3-319-05431-5_3-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000548_978-3-319-05431-5_3-Figure4-1.png",
        "caption": "Fig. 4 The leg orthosis of the MotionMaker",
        "texts": [
            " Both position and force control are taken into consideration. Muscles are electrostimulated in closed loop to reach the desired forces at the feet. The recruited muscles are presented and the clinical trials are discussed. 3 The MotionMaker is a product commercialized by Swortec SA. The MotionMaker is a mobilization robotic device for lower limbs. This section introduces the MotionMaker device, its components and explains how the therapy combining motion and electrostimulation works. The MotionMaker is based on two leg orthosis (Fig. 4). Thanks to screw based transmission and a crank mechanism, each leg orthosis allows the mobilization of the hip, the knee and the ankle joints. Each joint is instrumented with a force sensor at the extremity of the screw for safety and control purposes. The subject is first transferred to the device by using a transfer table (Fig. 5) and is then adjusted in a sitting position thanks to different motors (vertical adjustment, pelvic adjustment and inclination of the back). The anthropomorphic lengths are also adjusted with respect to the subject"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001025_2015-01-1477-Figure37-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001025_2015-01-1477-Figure37-1.png",
        "caption": "Figure 37. Scratch mark location and orientation from roll 3-5/8 ground contact (top) and plan view depicting scratch marks with velocity vector during ground contact (bottom).",
        "texts": [
            " As the component of perimeter velocity due to roll approaches the translational speed of the vehicle, the relative contribution of the direction of the CG velocity is decreased and the contribution resulting from roll velocity is increased. Additionally, the small effects of pitch and yaw rate on perimeter velocity will become more significant when the roll rate contribution approaches the CG velocity. For the left fender ground contacts, the perimeter speeds due to roll velocity were determined to be within 4 mph of the CG velocity during the 3rd and 4th revolutions. This condition resulted in abrasion patterns that were not parallel to the orientation of the CG velocity vector. Figure 36 and Figure 37 illustrate the scratch marks resulting from the roll 2-5/8 and roll 3-5/8 ground impacts located on the left front fender, hood, and left A-pillar. Scratch marks are first shown projected onto a CAD model of the pickup in the top portion of the figures. The lower illustration depicts the left front fender and hood with portions of the CAD model cut away. The model was oriented in the position at the ground contact, and the vehicle velocity vector is shown by the orange arrow. The scratch marks can be observed to be at a greater angle than the velocity vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003420_j.matpr.2020.07.486-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003420_j.matpr.2020.07.486-Figure5-1.png",
        "caption": "Fig. 5. Contact stress contour plot [9].",
        "texts": [
            " Kmax and Dk are not independent to each other; also define the load condition based on any two variables. The experimental independent geometry and material properties are expressed by da/dN and Dk, and considered the environmental and loading conditions The FEA is a familiar numerical modeling method and is used to analyze complex and non-linear behaviors of components. The commercial software ANSYS Workbench.12.1 used as a FEA tool in this analyzation work, that model is shown in Fig. 4. The contact zone had changed into elliptical area at after loading condition that shown in Fig. 5, and the reduced model consists of half of the elliptical area as shown in the figure. Also the contact pressure to the each load cases are extracted and plotted in the Fig. 6, and it shows that the relationship of analytical and FEA developed model of contact pressure in the large diameter bearing. The fatigue tool is used Please cite this article as: R. Pandiyarajan, K. Arumugam, M. P. Prabakaran et al., Fatigue life and fatigue crack growth rate analysis of high strength low alloy steel (42CrMo4), Materials Today: Proceedings, https://doi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000129_ecce.2019.8911875-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000129_ecce.2019.8911875-Figure1-1.png",
        "caption": "Fig. 1. IPMSMs with flat- and V-shape configurations.",
        "texts": [
            " From the observations of the calculated results, a slit in the rotor is proposed to reduce the cross magnetization. After confirming the validity of the FEA by experiments of a prototype motor, the slit shape is determined by automatic shape optimization. Finally, the characteristics of the optimized motor are compared with those of existing motors II. EFFECT OF CROSS MAGNETIZATION First, the effects of the cross magnetization are investigated by analyzing the torque components of the IPMSMs. Fig. 1 shows the motors with flat- and V-shape PM configurations. Table I lists the specifications. In both the motors, sintered Nd-Fe-B magnets are applied and the total magnet volumes are set to be identical to each other. One magnet is subdivided into 7 pieces along the axial length to prevent the eddy currents. The stator has 24 slots with concentrated windings. Both the stator and rotor cores are laminated. The inverter is an insulated gate bipolar transistor type PWM inverter whose carrier frequency is 10 kHz"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001548_iciea.2015.7334299-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001548_iciea.2015.7334299-Figure7-1.png",
        "caption": "Fig. 7. Grid subdivision",
        "texts": [
            " Equation(11) shows eddy current loss is proportional to the square of alternating frequency and the square of maximum flux density , is inverse proportional to the resistivity , and related to the structural parameters of PM. B. Simulation model A Two-dimensional model of FSCW-PMSM with 24-slot 16-pole is built by using FEM. On the basis of cycle symmetry, the whole model is divided into 1/8, and it is shown in Fig.6. Considering different area of the motor and the influence of skin effect, the grid subdivision shown in Fig.7. Fig.8 and Fig.9 represent the distribution of the flux density and the magnetic field lines on the motor model. The PM eddy current loss under rated-load condition is shown in Fig.10. It fluctuates periodically, because it changes with the change of the stator coil position. 1248 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) C. Eddy current loss of main harmonics The fundamental harmonic of air gap field has the same speed as rotor, so it does not produce eddy current on PMs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003478_icra40945.2020.9196890-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003478_icra40945.2020.9196890-Figure3-1.png",
        "caption": "Fig. 3: Structural Sketch of manipulator",
        "texts": [
            " Kinematic modeling of manipulator is the basis of trajectory planning and real-time control. The forward kinematic model of the manipulator must be accurate, and the inverse kinematic solutions can be obtained quickly and ensure the continuity of trajectory. The KC-30 robot has eight joints, of which the 3rd and 6th joints are expansion joints and the other joints are revolving joints. In this paper, kinematics model of 8-DOF manipulator is built based on D-H method. The reference coordinates of all joints are shown in Fig 3. Three virtual joints are designed in the model of forward kinematics to ensure that model is bulit along fuselage. Thus the measurements required by the model can be obtained from the CAD drawings. Based on Fig.3, D-H parameter table is shown in table 1. TABLE I: D-H Parameter Table Joint \u0398i di(mm) ai(mm) \u03b1i Offset 1 \u03981 770 0 pi/2 0 2 \u03982 0 557 pi/2 pi/2 Virtual 1 0 2670 0 0 0 3 0 L3+1030 0 -pi/2 0 4 \u03984 0 485 pi/2 0 Virtual 2 0 1800 0 pi/2 pi/2 5 \u03985 1425 0 -pi/2 0 Virtual 3 0 16.359 0 0 0 6 0 L6+2693.7 0 0 0 7 \u03987 663 0 pi/2 0 8 \u03988 1138 464 pi/2 pi/2 The coordinate transformation matrix between coordinate 7091 Authorized licensed use limited to: Dalhousie University. Downloaded on September 19,2020 at 07:19:44 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003990_978-3-319-14532-7_47-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003990_978-3-319-14532-7_47-Figure2-1.png",
        "caption": "Fig. 2 Strain gauges mounted in each pin hole of the test carrier planets",
        "texts": [
            " An arm is fixed on this ring and allowing the introduction of external load ((Hammami et al 2015) and (Hammami et al 2014a)). The two planetary gears are connected back-to-back in order to provide a mechanical power circulation: the sun gears of both planetary gear sets are connected through a common shaft and the carriers of both planetary gear sets are connected to each other through a rigid hollow shaft (Hammami et al 2014b). In order to compare the load sharing between the tests planets, three strain gauges are installed in the pin holes of each planet in the tangential direction of the test carrier (Fig. 2). The wires from the strain gauges are connected to the acquisition system through a hollow slip ring (HBM SK5/95) which is installed with the hollow shaft that connects the carriers (Fig.3). Strains gauges are used in quarter bridge configuration. Strain signals registered by strain gauges will be acquired by Programmable Quad Bridge Amplifier module (PQBA) of the acquisition system \u201cLMS SCADAS 316 system\u201d. This module can support four channels of strain transducers, piezo-resistive or variable capacitor sensors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003055_s12289-020-01560-1-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003055_s12289-020-01560-1-Figure10-1.png",
        "caption": "Fig. 10 Mean dimensional tolerances of all samples by FEA results",
        "texts": [
            " Samples may be deformed because of the residual stresses. Samples can be affected by the stress relaxation phenomena. Firstly, dimensional tolerances of all samples were measured at two heights. The first measurements were close to the building platform, while the second one was close to the top surfaces of samples. Both the top and bottom areas of samples were examined. Dimensional tolerances of samples at two levels are shown in Fig. 9 using the experimental measurements. Dimensional tolerances of samples at two levels are shown in Fig. 10 by FEA analysis (Autodesk Netfabb Simulation Utility). According to ISO 286, IT tolerances of samples were added. Bottom areas of samples deformed less than upper areas. Apparently, the height of samples has considerable influence on displacements. The building platform plays a fixation role, especially for the bottom areas. Bottom areas were tried to deform similarly as upper areas. However, the building platform kept those areas stationary. Since bottom areas were not deformed, internal stresses accumulated at the areas close to the building platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003355_ccdc49329.2020.9164756-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003355_ccdc49329.2020.9164756-Figure1-1.png",
        "caption": "Figure 1: Model of industrial robot",
        "texts": [
            " MDH model and the error model are established firstly. Considering the position deviation of the end target ball, its position is also taken as the parameter to be estimated, and then the parameters are identified by the LM method. Considering the measurement error, the previous result is taken as the initial value, and the parameters are further optimized by the PF algorithm. Finally, experimental verification and comparison are carried out. The universal serial 6-DOF industrial robot model is shown in figure 1. Using the DH method to model the 6-DOF in- dustrial robot, and the transformation matrix of its adjacent joints is Ti=Rot(z, \u03b8i)Trans(0, 0, di)Trans(ai, 0, 0)Rot(x, \u03b1i) (1) Expressed by the homogeneous transformation matrix, that is Ti= \u23a1 \u23a2\u23a2\u23a3 cos(\u03b8i)\u2212 cos(\u03b1i) sin(\u03b8i) sin(\u03b1i) sin(\u03b8i) ai cos(\u03b8i) sin(\u03b8i) cos(\u03b1i) cos(\u03b8i) \u2212 sin(\u03b1i) cos(\u03b8i)ai sin(\u03b8i) 0 sin(\u03b1i) cos(\u03b1i) di 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6 (2) Where, di is the offset of connecting rod, \u03b8i is the joint Angle, ai is the length of connecting rod, and \u03b1i is the included angle of joint axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000049_s00170-019-04441-3-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000049_s00170-019-04441-3-Figure3-1.png",
        "caption": "Fig. 3 Numerical domains and boundary conditions",
        "texts": [
            " In addition, a light source was set up beside the coaxial nozzle and projected through the falling powder in order to improve the visualization of the powder-gas flow. Finally, the images were taken against a black background in order to improve the resolution of the observation results. The nozzle head contained four radially-symmetric powder inlet jets orientated at an angle of 60\u00b0 relative to the horizontal axis and a laser protective gas jet. Due to the symmetry of the nozzle head, the simulations considered only a quarter-model, as shown in Fig. 3. An inlet velocity boundary condition was set at both the inlet of the protective gas nozzle and the inlet of the powder nozzle. In addition, a wall boundary condition was set on the surrounding walls of the powder nozzle and protective gas nozzle. A pressure outlet boundary condition was set on the cylindrical zone covering the volume beneath the nozzle [1]. Table 1 summarizes the powder and processing parameters considered in the DED simulations. In accordance with the specification of the experimental nozzle head, the diameter of the powder inlet was set as 3 mm and the diameter of the protective gas flow was set as 10 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003291_012008-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003291_012008-Figure3-1.png",
        "caption": "Figure 3. The surface roughness described in length",
        "texts": [],
        "surrounding_texts": [
            "Generally, when the vehicle moves on the road, the forces acting on the axle are dynamic and they are determined by solving the differential equation system describing the general dynamic model of the vehicle. In this study, the author limited the testing of vertical Fz31 and Fz32 reaction tests from the road surface to the axles. To verify and compare with the theoretical model developed, the author uses the input agitation function, which is the sinusoidal scale of the road surface. The height of the sinusoidal model is determined by the following formula: 1 1 cos 2 0 2( ) 0 0 , x H when x L Lh x when x x L                   H - maximum surface roughness height, H= 0.1m; L - surface roughness length (L= 0.6m); v - velocity of vehicle movement."
        ]
    },
    {
        "image_filename": "designv11_30_0003511_j.triboint.2020.106696-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003511_j.triboint.2020.106696-Figure4-1.png",
        "caption": "Fig. 4. Schematic diagram of friction and wear test rig.",
        "texts": [
            " Tribology International 154 (2021) 106696 where \u03a9 is rotating speed, rpm. d, l, \u03b2 are directly determined by geometrical structure. E is a common physical parameter and can be looked up in a material manual. KHb and KHc are also physical parameters, but the published data associated with it is not very abundant. Our group has developed a simple and effective method to measure KHb and KHc for a specific tribopair. The test method is briefly introduced in this article. The schematic diagram of the test rig is shown in Fig. 4. Take KHb as an example to introduce the measuring method. First of all, plate the ball sample with the same coating as the surface of the rotor, and make the disc sample of bristle material. The wear experiment worked with preset normal load, rotating speed and time interval. After the experiment, a scar left on the surface of the disc sample, the section area of which can be measured by 3D morphometer. KHb can be inferred with the preset and measured data based on equation (3). KHb = Volb FNL (30) The basic contents in tribology of the friction and wear test are shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000478_s12206-015-0703-z-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000478_s12206-015-0703-z-Figure2-1.png",
        "caption": "Fig. 2. Compliant shim structure.",
        "texts": [
            " The deformation of the shims in this kind of configuration can be of the same magnitude as the gas film thickness. In case of shock and vibration, the compliant shim can provide necessary deformation space for the pads. Meanwhile, the elasticity of the shim can be designed elaborately to provide variable stiffness for the tilting pads in radial direction. To approach this, the bolt at one end of the shim reserves a small distance vr as shown in Fig. 1(b). The geometrical character of the shim is shown in Fig. 2. The small circle at left is for the fastening bolt and the long and narrow hole at the right side is reserved for the displacement restriction bolt. The bigger circle of diameter fDh is for *Corresponding author. Tel.: +86 29 82664921 E-mail address: laitianwei@mail.xjtu.edu.cn \u2020 Recommended by Associate Editor Eung-Soo Shin \u00a9 KSME & Springer 2015 assembling with the pivot pedestal. The effective diameter of forcing on the shim is fDf, namely, the diameter of the pedestal. For the shim, it is a two-dimensional plate deformation problem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure10-1.png",
        "caption": "Fig. 10 Equal Force Wrench Sketch of the Segment",
        "texts": [
            " f j k is the force vector which is caused by the tension applied at the hole j on the upper plate of the segment k. The force vector f j k and f j k can be determined as: f j k = fke j k , f j k = \u2212fke j k+1 (12) Where, ej k = d j k\u2225\u2225\u2225d j k \u2225\u2225\u2225 , ej k+1 = d j k+1\u2225\u2225\u2225d j k+1 \u2225\u2225\u2225 . Substituting (4) and (12)-(11), the virtual work equation can be simplified as: f T\u03b4l = \u03c4T\u03b4\u03b8 (13) Where, f = [ f1 f2 \u00b7 \u00b7 \u00b7 f15 ]T , \u03c4 = [ \u03c41 \u03c42 \u00b7 \u00b7 \u00b7 \u03c410 ]T . From the Eq. 8, the variation of rope length can be calculated: \u03b4l = Jl\u03b4q (14) Substituting the Eqs. 14\u201313, the virtual torque can be represented as: \u03c4 = Jl Tf (15) In Fig. 10, when j = k, N + k, 2N + k, the rope j is the driving rope of the segment k and caused the equal force wrench S j k at the segment k: S j k = \u23a1 \u23a3 f j k \u03c1 j k \u00d7 f j k \u23a4 \u23a6 = [ fke j i fk \u03c1 j k \u00d7 e j k ] = [ e j k \u03c1 j k \u00d7 e j k ] fk = h j kfk (16) where, uj k = [ e j k \u2212 e j k+1 \u03c1 j k \u00d7 e j k \u2212 \u03c1 j k \u00d7 e j k+1 ] , hj k = [ e j k \u03c1 j k \u00d7 e j k ] . In Fig. 10, when j = k+1 \u00b7 \u00b7 \u00b7N, N +k+1 \u00b7 \u00b7 \u00b7 2N, 2N + k + 1 \u00b7 \u00b7 \u00b7 3N , the equal force wrench of the segment k can be represented as: S j k = \u23a1 \u23a3 f j k + f j k \u03c1 j k \u00d7 f j i + \u03c1 j k \u00d7 f j k \u23a4 \u23a6 = [ fke j k \u2212 fke j k+1 fk \u03c1 j k \u00d7 e j k \u2212 fk \u03c1 j k \u00d7 e j k+1 ] = u j kfk (17) Where, uj k = [ e j k \u2212 e j k+1 \u03c1 j k \u00d7 e j k \u2212 \u03c1 j k \u00d7 e j k+1 ] According to Eqs. 16 and 17, the equal force wrench of the segment k can be represented as: Sk = J t kf (18) Where, J t k is the tension-wrench Jacobin matrix of the segment k, and can be calculated: J t k = \u23a1 \u23a2\u23a3 06\u00d7(N\u22121) h j k u j+1 k u j+2 k \u00b7 \u00b7 \u00b7 uN k 06\u00d7(N\u22121) h j+N k u j+N+1 k u j+N+2 k \u00b7 \u00b7 \u00b7 u j+2N k 06\u00d7(N\u22121) h j+2N k u j+2N+1 k u j+2N+2 k \u00b7 \u00b7 \u00b7 u j+3N k \u23a4 \u23a5\u23a6 (19) The equal torque can be calculated from the equal force wrench Sk is: \u03c4 = N\u2211 k=1 J q k Sk (20) Where, J q k = \u23a1 \u23a2\u23a3 z0 (p2k \u2212 p0) \u00d7 z0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002561_iros.2016.7759080-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002561_iros.2016.7759080-Figure3-1.png",
        "caption": "Fig. 3. Serial Hybrid chain model of musculoskeletal dual arm manipulator",
        "texts": [
            " This musculoskeletal structure can be classified as a general hybrid-type closed-chain mechanism, since several closed-chains are embedded in the serial structure. Contribution of muscles to impulse will be analyzed in comparison to arm model without muscle. The external impulse given by (8) and the velocity vector of colliding bodies are assumed to be known. Fig. 2 describes the flowchart of the impulse modeling for the musculoskeletal dual arm model. To find the impulses at cut-joints, the dual closed-chain mechanism is transformed into two serial hybrid systems as shown in Fig. 3 which is created by cutting one joint. The two serial hybrid structures can be achieved by cutting ( 3  ) or ( 6  ), the results are same. The cut-joint h  is shown in Fig. 3. The impulses 1 2 \u02c6 \u02c6, h h F F at cut-joint act with same magnitude but opposite in direction. The contact point with the environment is located on left chain. Refer to Imran and Yi [17] for the dynamic model of each serial hybrid chain structure the dynamic model for each serial hybrid chain system is given as follows [ ] [ ] [ ] [ ] , T h T T I T T i i i i i i i i h i extI P G G      T F F   (9) where (i 1,2), [ ] I i G and [ ] h i G are the first-order kinematic influence coefficient (KIC), relating the output velocity vector I  and the cut-joint velocity vector h  to the joint velocity vector i  for each serial hybrid chain, respectively",
            " The velocity increment at cut-joint can be described with respect to the velocity increment at i  Now integrating (9) over the contact time, we have 1 1\u02c6 \u02c6[ ] [ ] [ ] [ ] .h T I T i i i i h i i ext I G I G       F F (11) Inserting (11) into (10) yields 1 1\u02c6 \u02c6 , T T h h h I i h i i i i h i i i extG I G F G I G F                                  \u02c6 \u02c6 . i h i h i h ext S F T F  (12) It is noted that the cut-joint velocity increment can also be describe in terms of the independent joint velocity as ,i i a aG      (13) where the independent joints of the dual arm system are denoted as three joints (e.g., 1 2 3 , ,   Fig. 3) since the mobility of the system is three. Now inserting (13) into (10) h h i h i i i i a aG G G                   . (14) Inserting (2) into (14), the final form of (14) becomes    \u02c6 \u02c6 . TTh I i h i i a aa a ext i h ext G G I G F U F                (15) By equating (12) and (15), the closed-form relationship between cut-joints and external impulses can be written as      \u02c6 \u02c6 \u02c6 . i h ext i h i h i h ext U F S F T F  (16) Finally, the impulses at cut joints are given by \u02c6 \u02c6 ,h i h i e extF M F    (17) where     1[ ] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001743_iccas.2015.7364719-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001743_iccas.2015.7364719-Figure3-1.png",
        "caption": "Fig. 3 Coordinate system ofHAUY.",
        "texts": [],
        "surrounding_texts": [
            "In order to control the depth of HAUV underwater, a depth controller is designed applying PID controller. Fig. 4 shows block diagram of the depth controller. The depth controller compares aimed depth( zd) to present depth( Z), and utilizing this data, it controls PID. The result value of control determines the power( Z) against the axis ofHAUV."
        ]
    },
    {
        "image_filename": "designv11_30_0001285_j.proeng.2014.06.068-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001285_j.proeng.2014.06.068-Figure1-1.png",
        "caption": "Fig. 1. Table tennis racket (a), blade (b), blade plywood composition (c)",
        "texts": [],
        "surrounding_texts": [
            "The perception of the table tennis racket performances depends on many factors. Most of the appreciations expressed by confirmed players are subjective since they are relative to their proper sensory analysis. Usually, the rackets can be qualified as fast, slow, stiff, adhesive, controllable, etc. It can happen that two players of same level will give contradictory opinion for the same racket. The acoustic signature at ball impact is the first appreciation element of the racket performances; therefore it has to sound well in order give a positive a priori for the other evaluation parameters. The ball-racket impact in table tennis had been studied in many works from the restitution coefficient point of view, the main results and ideas can be found in Kawazoe et al. (2003) and in Tienfenbacher et al. (1994, 1996). To our knowledge, there is no published work on the vibro-acoustic of tennis table racket. Nevertheless, numerous informations can be found and read on the internet forums and blogs edited by players and competitors. This work is in the continuity of previous studies that had analyzed the vibratory behavior of racket blades and also the vibro-acoustic following the ball impact, Manin et al. (2012). The vibration modes were simulated and correlated satisfactorily with some experiments. The study presented here is concerned with the vibro-acoustic of table tennis racket at ball impact and more specifically on the influence of the racket blade plywood composition: type of woods, thickness of the plies. First, some previous work results are given about the correlation that exists between the structure and the acoustic vibrations. The structure vibration modes that produced the sound at ball impact were identified. Then, several racket blade prototypes have been realized to study the influence of the racket blade plywood composition on the vibro-acoustic behavior. The blade plywoods tested differed by the thickness of the plies and/or their wood essence. In parallel, a sensory analysis had been conducted with some high level players. Finally, the sensations of the players are analyzed versus the laboratory test results obtained."
        ]
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure1.8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure1.8-1.png",
        "caption": "Fig. 1.8 Structure of the moving-coil-type linear DC motor [22]",
        "texts": [
            "3 Linear machines for applications of logistics [11]. . . . . . . . . 4 Figure 1.4 Linear machines for applications of Industrial Automation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 Figure 1.5 Linear machines for applications of rail transportation . . . . . 6 Figure 1.6 Winding arrangements for tubular PM linear machines [19] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 Figure 1.7 An axial flux permanent magnet motor [21] . . . . . . . . . . . . 9 Figure 1.8 Structure of the moving-coil-type linear DC motor [22] . . . . 10 Figure 1.9 Structure of the moving-coil-type linear DC motor [23] . . . . 10 Figure 1.10 Schematic of reciprocating linear generator system [24] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.11 A linear vibration-driven electromagnetic micro-power generator [26] . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.12 An improved axially magnetized tubular PM machine topology [28] ",
            " A double-sided slotted torus axial flux permanent magnet motor suitable for direct drive of electric vehicle is proposed in [21]. Its construction is illustrated in Fig. 1.7. The magnet poles are arranged alternatively along the rotor circumference, and magnetized in axial direction. The motor can be easily mounted compactly onto a vehicle wheel, fitting the wheel rim perfectly. 1.4 Magnet Patterns of Linear Machines 9 Nirei et al. presents the evaluation of a moving-coil-type cylindrical linear DC motor [22]. The structure is illustrated in Fig. 1.8. It consists of a coil, permanent magnets, yokes, a coil bobbin held by the arms of a coil holder and linear bearings. There are 16 pieces of permanent magnets, and they are fitted along the inner surface of the outer yoke and magnetized radially. The yokes are solid steel blocks. The inner yokes are separated into four pieces to provide space for the arms of the coil holder. Kim et al. presents a tubular linear brushless permanent magnet motor as shown in Fig. 1.9 [23]. It has a slotless stator to provide smooth translation without cogging"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-10-1.png",
        "caption": "Figure 6-10. Step 3 to obtain the propulsion or allocation matrix",
        "texts": [
            " Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 253 Step 3: Choose the direction of rotation of the motors. Remember that in order to not produce a self-rotating effect, half of the drone\u2019s motors must turn in a clockwise direction and the other half in a counterclockwise direction. (When dealing with a vehicle with oddnumbered motors, one of them is used as a vectorizer. See the appendix on thrust vectoring). In general, there are several ways to achieve this, but the most usual is a diagonal sequence (see Figure\u00a06-10). There are efficiency studies on which motors must have a certain direction of rotation. Moreover, you are free to experiment. However, in this example and for a certain degree of standardization we will choose the direction of rotation shown in Figure\u00a06-10. Step 4: Choose the geometric configuration. In this case, place the arms at the same distance to the center of the vehicle, with square angles (0 and 90 degrees) and at the same height with respect to the center of the vehicle. Very unequal distances, very different angles, or significantly different heights change the propulsion matrix considerably. Nevertheless, if the differences are moderate, this does not influence the operation of the vehicle. See Figure\u00a06-11. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 254 Step 5: Place the coordinate frame of the drone (that is the body frame) and label the axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003574_2374068x.2020.1835007-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003574_2374068x.2020.1835007-Figure1-1.png",
        "caption": "Figure 1. Methodology of present work.",
        "texts": [
            " obtained through FDM and VS can be further used as sacrificial patterns in IC applications. But hitherto little has been reported on combining FDM, VS and IC for controlling surface roughness and dimensional accuracy of biomedical implants. In this paper, it is proposed to investment cast biomedical implants with intermediate VS of FDM patterns and the process parameters of these three processes are optimised to get the best surface finish and dimensional accuracy of hip joint (selected as biomedical implant). The various steps of the study are illustrated in Figure 1 in the form of a flowchart. A biomedical implant such as hip joint used in human body has been identified as benchmark. Stainless steel 316 L has been used to cast these parts through the rapid IC process. The drawing of implant has been shown in Figure 2. This work is aimed in a bid to produce the benchmark with improved surface finish. The various input factors that may influence the dimensional accuracy and surface finish have been identified. These factors are: orientation of pattern, density of pattern, no"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-9-1.png",
        "caption": "Figure 6-9. Step 2 to obtain the propulsion or allocation matrrix",
        "texts": [
            " Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 252 Step 1: Establish the symmetry of the vehicle. This influences how to place the body frame, as well as how the motors influence certain movements. This is performed based on the guide-mark of the autopilot, which indicates the zero reference in the yaw angle from your IMU sensor or the magnetic north of the compass in other models. See Figure\u00a06-8. Step 2: Assign a number to each motor. This affects even the way in which the motors will be connected to the autopilot. See Figure\u00a06-9. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 253 Step 3: Choose the direction of rotation of the motors. Remember that in order to not produce a self-rotating effect, half of the drone\u2019s motors must turn in a clockwise direction and the other half in a counterclockwise direction. (When dealing with a vehicle with oddnumbered motors, one of them is used as a vectorizer. See the appendix on thrust vectoring). In general, there are several ways to achieve this, but the most usual is a diagonal sequence (see Figure\u00a06-10)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure2-1.png",
        "caption": "Fig. 2. Detail view of the fabricated camera module",
        "texts": [
            " Also it has a sonde system to locate a point where the PIBOT system runs in the buried pipeline. The camera module was designed to integrate computing, wireless communication, video-sensing, lighting and emergency-locator systems into a single monolithic module. And it consists of battery pack to provide power for the camera module, the main control board, the sub control board, the locating-sonde coil, the forward LEDs, charging contacts and the fisheye camera-dome. The camera module is shown in Fig. 2. The drive module includes the ability to keep the robot body in the center of pipeline with the driving wheels which have the support arms. All the required mechanical elements, electrical PCBs, various sensors and odometers were integrated by the design shown in Fig. 3. The module was designed to expand/collapse a set of four drive-legs, and each wheel system contains a harmonic gear to rotate a drive-wheel using In-wheel motor. The battery module provides 25.2 VDC and up to 70 Ah. of energy to the robot for a meaningful 10 h mission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002070_indicon.2015.7443345-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002070_indicon.2015.7443345-Figure1-1.png",
        "caption": "Fig. 1. Maglev system model",
        "texts": [
            " Section V, simulation results are presented to emphasize the efficiency of proposed This work was supported by Indian Council for Cultural Relations (ICCR) under Africa Scholarship Scheme, under the program of executive program between Arab Republic of Egypt and India, with ICCR Ref. No. ASS-248\\2014-2015, Ministry of External Affairs of India. 978-1-4673-6540-6/15/$31.00 \u00a92015 IEEE methods and compare it with PDC scheme. Finally in section VI , the conclusion is given. Notation: The superscript \u201cT\u201d represents the transpose of a matrix. The notation \u201c*\u201d is used as an ellipsis for terms that are induced by symmetry. II. MAGLEV DYNAMICS Fig.1 shows Magnetic Levitation (Maglev) system model. The control purpose is to stabilize the position of the ball by the force of electromagnet. The Maglev's dynamics equations [7] are presented as 2 \u239f \u23a0 \u239e \u239c \u239d \u239b= +=+ h ikf fmgfhm d (1) Where m is the mass of the iron ball ; h is the position of ball; fd is the disturbance force; g is the acceleration gravity; k is the electromagnet constant; i is controlled current of the electromagnet. Define the state variable as of (1) as: hx,hx == 21 Then the state equation dfmx i m kgx xx 1 2 1 2 21 +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u2212= = (2) III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000745_s1560354714010079-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000745_s1560354714010079-Figure5-1.png",
        "caption": "Fig. 5",
        "texts": [
            " 1 2014 We direct the x-axis along the rod, and the y-axis lies within the base plane and contains the rod center (at the initial time). Any future position will be defined by the abscissa of the rod center q1 and by the ordinates of the particles q2 and q3. The kinetic energy matrix at the initial position is: A = diag {2m,m,m} while the external forces can be represented as the sum of three forces: Xis directed along the rod, and Y1 and Y2 are applied to the particles orthogonally to the rod (Fig. 5). The friction forces have components (Tx1, Ty1) and (Tx2, Ty2), which satisfy the following inequalities: T 2 xj + T 2 yj P 2, P = \u03bcmg, j = 1, 2, (2.14) where \u03bc is the coefficient of friction for both particles. According to Theorem 2, these forces can be found from minimality conditions for the function \u03a6(T ) = 1 2m (X + Tx1 + Tx2) 2 + 1 m (Y1 + Ty1) 2 + 1 m (Y2 + Ty2) 2 (2.15) on the set (2.14). There are the following four possibilities: 1) The minimum of the function (2.15) is zero (the system is at rest)",
            "21) where the signs of the radicals are chosen following the idea of minimizing the expression (2.15). System (2.21) reduces to the single equation u = H2 (H1(u)) , u \u2208 (\u22121, 1), (2.22) then the value v can be calculated from the second formula (2.21). If Eq. (2.22) has several roots, we choose the root, where the value (2.15) is the smallest. For an illustration of such an approach we suppose that the external forces are equivalent to the single force of value S acting along the line which passes at distance d from the origin and forms angle \u03b1 with the rod (Fig. 5). Substitution into (2.15) gives X = S cos \u03b1, Y1 = 1 2 S(sin \u03b1 \u2212 \u03ba), Y2 = 1 2 S(sin \u03b1 + \u03ba), \u03ba = d l . Set P = 1, d = 2l, \u03b1 = \u03c0/6 and increase S gradually from zero. Within the interval 0 < S < S\u2217 = 0.8 inequality (2.16) is valid, hence, the system is at rest. For S = S\u2217 the second radical in (2.16) vanishes; the interval S\u2217 < S < S\u2217\u2217 = 4/ \u221a 21 \u2248 0.873 corresponds to the rotation around the first particle in accordance with (2.18). For S = S\u2217\u2217 the first expression in (2.17) turns into an equality, and for S > S\u2217\u2217 the rod start sliding at both contact points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000601_crv.2015.46-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000601_crv.2015.46-Figure6-1.png",
        "caption": "Figure 6. Quadcopter reference frame used for simulation [24]",
        "texts": [
            " A cube shaped obstacle with 500 randomly distributed points is generated while the entire control architecture for the collision avoidance system discussed in Section IV is modeled on a quadcopter Unmanned Aerial Vehicle (UAV). An open 1Since we simulate this system on a quadcopter UAV, the maximum control input in the z direction is the maximum available thrust while the maximum control input in the x and y directions corresponds to the maximum allowable roll and pitch angles, respectively. source quadcopter dynamics controller is used to simulate vehicle dynamics. The simulation reference frame is shown in Figure 6. The following values are used for the various parameter of the system: 1) A maximum deceleration (amax) of 5m/s2 is chosen for this simulation. This corresponds to a maximum lateral g-force of 0.5g. 2) The radius of sphere of influence (rsoi) is set to 5m. 3) The constant of proportionality (\u03bc) is set to 20. This corresponds to an increase or decrease in the magnitude of the potential field which corresponds to aggressiveness in the vehicle response. 4) The maximum force in the x, y, and z directions is set at 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003055_s12289-020-01560-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003055_s12289-020-01560-1-Figure1-1.png",
        "caption": "Fig. 1 Configurations of three manufacturing sets (a) one sample, (b) five samples (c) thirteen samples",
        "texts": [
            " Material properties were defined according to information that was given by the suppliers [30]. In this study, three sets of samples were manufactured. All samples were cubic and had the same dimensions (10x10x10 mm). The first set contained only one sample on the manufacturing platform, while the second manufacturing contained five samples. Samples positioned symmetrically with 5 mm space between each other. Finally, the third set of manufacturing contained thirteen samples on the platform. Samples were positioned symmetrically with a 5 mm distance. Figure 1 shows the placement of manufacturing sets. Thermal histories of all manufacturing processes were recorded via a thermal camera. Optris PI160 thermal imager was used. The camera resolution is approximately 1 mm, and the measurement speed is 7.8 \u00d7 10\u22123 fps (frames per second). The measurable temperature values by the camera are ranged between - 100 \u00b0C and 1500 \u00b0C. The thermal camera was mounted inside the isolated manufacturing chamber of the machine. The temperature distribution of the processes was recorded with 350 mm distance from samples"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000228_j.actaastro.2019.12.021-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000228_j.actaastro.2019.12.021-Figure6-1.png",
        "caption": "Fig. 6. Force analysis of upper part of each RPS chain.",
        "texts": [
            " (34) The Newton\u2212Euler equation of the rigid body can be written as292 3\u2211 i=1 fui +msg + fd = msk\u0308, (35) 3\u2211 i=1 vi \u00d7 fui + nd = Is\u03c9\u0307 + \u03c9 \u00d7 (Is\u03c9) , (36) where ms is the mass of the rigid body; Is is the moment of inertia of the rigid293 body about its mass center in O\u2212XY Z; Is = RI \u2032sR T, and I \u2032s is the moment294 of inertia of the rigid body about its mass center in o\u2212 xyz; fd and nd are295 the external force and moment imposing on the rigid body, respectively.296 Due to Jvi = [vi\u00d7] =  0 \u2212viz viy viz 0 vix \u2212viy vix 0 , Eqs. (35) and (36) can be297 rewritten as298 3\u2211 i=1 fui = msk\u0308 \u2212msg \u2212 fd, (37) 3\u2211 i=1 Jvifui = Is\u03c9\u0307 + \u03c9 \u00d7 (Is\u03c9)\u2212 nd. (38) Combining Eqs. (29), (37), and (38) yields299  1 1 1 Jv1 Jv2 Jv3 Jf1 0 0 0 Jf2 0 0 0 Jf3   fu1 fu2 fu3  =  msk\u0308 \u2212msg \u2212 fd Is\u03c9\u0307 + \u03c9 \u00d7 (Is\u03c9)\u2212 nd Q1 Q2 Q3  . (39) 17 Eq. (39) can be rewritten as300 Jzfu = Q. (40) Based on Eq. (40), fu can be computed by301 According to Fig. 6, based on Newton\u2019s equation, the force balance e-303 quation of the upper part of each RPS chain in Ri \u2212 xiyizi can be expressed304 as305  RfPi,xi + Rfui,xi = m2 Rv\u0307ui,xi RfPi,yi \u2212 Rfui,yi \u2212m2 Rgyi = 0 RFi \u2212 Rfi \u2212 Rfui,zi \u2212m2 Rgzi = m2 Rv\u0307ui,zi , (42) where Rfui = RT i fui; Rv\u0307ui = RT i v\u0307ui; Rgyi = m2gc sin (\u03b8i); Rgzi = m2gc cos (\u03b8i);306 gc is the gravitational acceleration.307 According to Eq. (42), we obtain308 18 { RfPi,xi = m2 Rv\u0307ui,xi \u2212 Rfui,xi RfPi,yi = Rfui,yi +m2 Rgyi . (43) The Coulomb friction force between the P pairs can be expressed as309 Rfi = \u2212\u00b5 \u221a (RfPi,xi )2 + (RfPi,yi) 2 \u00b7 sgn ( Rvui,zi ) , (44) where \u00b5 is the friction coefficient; Rvui = RT i vui; if Rvui,zi > 0, sgn ( Rvui,zi ) =310 1; if Rvui,zi = 0, sgn ( Rvui,zi ) = 0; if Rvui,zi < 0, sgn ( Rvui,zi ) = \u22121"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003558_j.sysconle.2020.104804-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003558_j.sysconle.2020.104804-Figure1-1.png",
        "caption": "Fig. 1. Cooperative cruise control methodology.",
        "texts": [
            " In this case, in the same way V\u0307 \u2264 0 if \u2212 N\u2211 i=1 \u03bbmin (Q ) N\u2211 j=1 \u2225eij\u22252 + 2 N\u2211 i=1 N\u2211 j=1 \u2225Pbi\u2225\u2225eij\u2225\u03f5\u2217 0 \u2264 0 \u21d2\u21d2 N\u2211 i=1 N\u2211 j=1 \u2225eij\u2225 \u2265 2\u2225Pbi\u2225\u03f5\u2217 0 \u03bbminQ , ith this condition we can then ensure that all synchronization rrors are bounded even when there is no communication of the nput between agents. \u25a0 5. Simulation results This section presents the study case of cooperative cruise control as an application of proposed Theorems 1\u20133 and the simulation results obtained. We introduce the cooperative cruise control, where a network of vehicles will maintain the same speed and distance between each other. The technique is known as cooperative adaptive cruise control. Fig. 1 shows the graphic interpretation of this methodology, where vi represents the speed present in each vehicle and di the distance between them. Each agent is modeled as x\u0307i = \u23a1\u23a3 0 1 0 0 0 1 0 0 \u2212 1 \u03c4i \u23a4\u23a6 xi + \u239b\u239d\u23a1\u23a3 0 0 1 \u03c4i \u23a4\u23a6+ fi(xi) \u239e\u23a0 ui, here \u03c4i represents the inertial time lag of the power-train sysem, that varies between each vehicle, according to its physical haracteristics and its environment. Input ui is the acceleration efined as the force per vehicle mass. In this case, there is a eading agent that defines the acceleration profile that the vehicle etwork must maintain"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure25-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure25-1.png",
        "caption": "Figure 25. Speed field contours for 1000 rpm.",
        "texts": [],
        "surrounding_texts": [
            "The first stage of the iterative algorithm for solving the problem of optimizing the kinematic and geometric characteristics of the field stripper and stationary installations has been developed. The task of studying the movement of the grain mass is divided into two stages. At the first stage, a model of air mass movement in the considered installations is built and based on a mathematical model that takes into account the turbulence of the movement using the finite volume method; the field of velocities and pressures is calculated. The influence of the technological parameters of the installations on these ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 fields was studied: the diameters and rotation speed of the stripping drum and the additional drum-fan, the shape of the surface of the upper deck of the installation chamber. The calculations have shown that additional field experiments are required to build an adequate model of the stationary installation. A series of calculations was carried out according to the developed algorithm; the geometric and kinematic parameters of the installations were obtained."
        ]
    },
    {
        "image_filename": "designv11_30_0001476_s12303-014-0049-z-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001476_s12303-014-0049-z-Figure4-1.png",
        "caption": "Fig. 4. Nomenclature and symbols used for the asymmetric boudin structural elements and geometric parameters, after Goscombe and Passchier (2003); parameters are defined in the text. Shear sense is dextral in this boudin train.",
        "texts": [
            " In this study geometry and kinematics of natural boudins are classified and determined, respectively, either directly in the field or through photographs. Measurement from photographs is considered as accurate as those measured in the field. The average absolute difference between measurements from outcrop and photograph of the same structure is on the order of \u00b16\u00b0 for angular parameters and 4\u20129 cm for longitudinal parameters. In this study we used a suite of nomenclature for structural elements and geometric parameters after Goscomb et al. (2004). These parameters can be observed in sections normal to boudin necks of individual boudin blocks (Fig. 4). Enveloping surface of boudinaged layer is equivalent to the overall trace of the boudin train. In most cases, enveloping surface of boudinaged layer is parallel to the penetrative foliation (Sp) in the host rock (Fig. 4). For any boudin type containing boudin faces or inter-boudin zones, an inter-boudin surface (Sib) can be defined. In asymmetric boudin structures, Sib is an inclined, discrete surface that coincides with the boudin face, along which boudins were laterally displaced and which arcs into parallelism with the enveloping surface (Se; Fig. 4). The inter-boudin surface develops between two distinct boundins, and is similar to a fault that separates these two terminating to the boudin exteriors which is the original surface of boudin block in contact with the host. A linear feature (Lb) can be described as the orientation of the boudin long axis, boudin edge or neck zone (Fig. 4). Lb is usually parallel to the intermediate axis (Y) of the finite strain ellipsoid. Sib and Lb can be highly oblique or have a perpendicular orientation relative to an extension axis (Le) and mineral stretching lineation. Lateral displacement (D) between individual boudins is measured along Sib and normal to Lb, whereas amount of dilation of Sib (N) or the width of the inter-boudin gap (Fig. 4) is measured normal to Sib (Goscombe and Passchier, 2003; Goscomb et al., 2004). In order to describe the geometry of the boudins, we have used L (length of individual boudin blocks) and W (width or boudin thickness). Both L and W are measured in the profile plane, so that L is measured parallel to Sb and W normal to Sb. \u03b8 (the angle between Sb and Sib), \u03b1 (relative block rotation or the acute angle between Se and Sb), and \u03b8\u2019 (the acute angle between Sib and Se) are also measured in the profile plane, where \u03b8\u2019 = \u03b8 \u2012 \u03b1 (Fig. 4). There are at least four different ways in which folds and boudins can be related in the field observations. (1) Fold and boudins may form synchronously as the result of a compression in one direction and an extension in the other direction. (2) Folds and boudins may form in sequence during a single deformation. (3) Boudins formed during one deformation may be folded during a later deformation and vice versa. (4) Depending upon the orientation of the maximum compression with respect to the layering or fabric, a complete range of structures can exists between symmetrical and asymmetrical folds, and symmetrical and asymmetrical boudins (Price and Cosgrove, 1990)",
            " Hyperbolic distribution method for estimating the kinematic vorticity number that is based on the premise that the orientation of the long axes of backward vergence of asymmetric boudins within the acute angle field between the flow eigenvectors delineates the orientation of the unstable eigenvector (Forte and Bailey, 2007). The stable eigenvector is assumed to be parallel with foliation (Simpson and De Paor, 1997). The axial ratio (R) of the boudins and the angle (\u03b1) between the long axis of boudin (L, Fig. 4) and the shear zone boundary or foliation were measured for about 220 samples in the XZ-plane, normal to the boudin axis (Lb, Fig. 4). Both \u03b1 angle and R values were plotted on a hyperbolic net (De Paor, 1988). The hyperbola represents the acute angle between the eigenvector orientations, where one eigenvector is fixed to the flow plane of the simple shear component and the orientation of the other (shortening) eigenvector depends on the pure shear contribution. One limb of this hyperbola is selected to be asymptotic to the mylonitic foliation, assuming that it is sub-parallel to the extensional flow apophysis, while the other is considered to delineate the orientation of the unstable flow apophysis (Xypolias, 2010)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure5-1.png",
        "caption": "Fig. 5. Planar substitution. (a) Common substitution: Pin joints at Cm and Cf are connected by an additional link, (b) general conjugate points K and Ko as substitution centers.",
        "texts": [
            " Then, \u03c9 l and \u03bc l can be uniquely solved for from the following equations: \u03c9 l r \u00bc vK=Ko \u00bc vK\u2212vKo \u00f02\u00de \u03c9 l \u03c9 l r\u00f0 \u00de \u00fe \u03bc l r \u00bc aK=Ko \u00bc aK\u2212aKo : \u00f03\u00de In addition, equations for velocity and acceleration of points K and Ko are obtained through constraining of the elements of the higher pair (for example, see the revolute joint constraints in Fig. 1). These set of equations can be solved for complete velocity and acceleration analysis of the mechanism containing this higher pair. The planar analogue of the virtual sphere is the osculating circle. In Fig. 5(a),m-curve and f-curve are the planar profile curves of the contacting rigid bodies in point contact. Cm and Cf are the centers of curvature ofm-curve and f-curve respectively at the point of contact C. For any relativemotionup to second-order, the center of curvature of the instantaneous point-path ofCm is located at Cf. This means that pointCm instantaneouslymoves on a circlewhose center is located atCf. This is an important property of planar enveloping curves [11]. The two centers of curvature can be joined by an additional linkwith revolute joints at the centers of curvature to give the classical substitute-connection",
            "We see that this substitution requiresmere inspection of the curves in contact. There can be other substitutions also. For this, any point of them-curve body can be chosen and its corresponding center of curvature of instantaneous pointpath is chosen in the f-curve body [8]. Let us find here the center of curvature of point-path of the chosen point belonging tom-profile body on the n-linewhich is not coincidentwith Cm. A point K is chosen on the n-line and be at a signed distance d from Cm as shown in Fig. 5(b). Point I be the location of the instantaneous center on the n-line and y be the signed distance of K from I. The square of the speed of point K, labeled v2, can be written as: v2 \u00bc \u03c92y2: \u00f04\u00de Similarly, the square of the speed of Cm, labeled vm 2 , is written as: v2m \u00bc \u03c92 y\u2212d\u00f0 \u00de2: \u00f05\u00de Let \u03c1m be the signed radius of curvature of instantaneous point-path of Cm. We have that \u03c1mj j \u00bc CmC f . Then the Y-component of acceleration of point Cm, labeled amy, can be written as: amy \u00bc v2m \u03c1m : \u00f06\u00de polar axis [10] of the curve V at a point K is a straight line parallel to the binormal and passing through the center of curvature corresponding to K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.23-1.png",
        "caption": "Fig. 18.23 Principle of ion mobility spectrometry",
        "texts": [
            " IMS both separates and detects chemicals, using differential migration of ions in an electric field. In IMS, gas phase species are required to be ionized, for example by high energy electrons from a radioactive 63Ni source. Ions travel in a gas stream through an electric deflection field that spatially separates the ions with respect to their ion mobility at atmospheric pressure. Ion species with different characteristics (mass, charge, and size) have different drift velocities, as in Eq. (18.13) where K is an ions mobility and vd is its drift velocity (see Fig. 18.23). Ideally, individual ion beams develop that are spatially separated. vd \u00bc KE \u00f018:13\u00de By increasing the deflection voltage of the electric field all ion beams are successively directed onto the collector electrode, where the ion current I is measured. Differentiating the recorded I(V ) curve results in the ion mobility spectrum. Under constant conditions the ion mobility K is a characteristic measure for a certain ion species. Typically an IMS will have a high resolution of R> 20: R \u00bc td=Wt, 1=2 \u00f018:14\u00de where td is the drift time andWt,1/2 the temporal peak width measured at half of the maximum peak height"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001113_10426914.2014.901531-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001113_10426914.2014.901531-Figure1-1.png",
        "caption": "FIGURE 1.\u2014The overlap configuration.",
        "texts": [
            " The objective of this investigation aims at the optimal process parameters on microstructure and mechanical properties in H62 brass to 316L stainless steel welding by Nd:YAG pulsed laser. The materials used in the pulsed laser experiments were 316L stainless steel and H62 brass. The thickness of 316L stainless steel and H62 brass sheets was 0.9mm and 0.4mm, respectively. The dimension of the 316L stainless steel sheets for the welding was 100mm 20mm 0.9mm (length width thickness). The dimension of the H62 brass sheets for theweldingwas 100mm 20mm 0.4mm (length width thickness). Brass-on-stainless steel overlap configuration was used in this study, as shown in Fig. 1. The most important thermophysical properties of the twomaterials are listed in Table 1. Apparently, the physical Received January 6, 2014; Accepted February 19, 2014 Address correspondence to Junqi Shen, Room C0804, Building 25, School of Materials Science and Engineering, Tianjin University, Weijin Road 92, Nankai District, Tianjin 300072, People\u2019s Republic of China; E-mail: shenjunqi@tju.edu.cn Color versions of one or more of the figures in the article can be found online at www.tandfonline"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003896_978-1-4471-5102-9_178-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003896_978-1-4471-5102-9_178-1-Figure2-1.png",
        "caption": "Fig. 2 Generalized coordinates for a single wheel",
        "texts": [
            " For example, both these vehicles require a specific device (differential) for distributing traction torque to the driving wheels. The starting point for modeling wheeled mobile robots is the single wheel. This may be represented as an upright disk rolling on the ground. Its configuration is described by three generalized coordinates: the Cartesian coordinates .x; y/ of the contact point with the ground, measured in a fixed reference frame, and the orientation of the disk plane with respect to the x axis (see Fig. 2). The configuration vector is therefore q D .x y /T . The pure rolling constraint is expressed as sin cos Px Py D 0 (1) and entails that, in the absence of slipping, the velocity of the contact point has a zero component in the direction orthogonal to the wheel plane. The angular speed of the wheel around the vertical axis is instead unconstrained. The kinematic constraint (1) is nonholonomic, i.e., it cannot be integrated to a geometric constraint; this may be easily shown using Frobenius theorem, a well-known differential geometry result on integrability of differential forms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure1-1.png",
        "caption": "Fig. 1 Imaginary generating gears and work gears: (a) pinion and (b) gear",
        "texts": [
            " This generating gear is a virtual gear whose teeth are formed by the locus of the cutter blades, although its tooth number is not necessarily an integer. In addition, whereas its rotation axis coincides with the rotation axis of the machine cradle, the rotation angles of the cradle and the work gear are relatively timed in the generating method but not in the nongenerating method. In this present study, both pinion and gear, being cut by the same planar-generating gear that results in conjugated tooth surfaces, are produced by the generating method (Fig. 1). Hence, if the cutters are arranged in the same position as the generating gear, they will produce a conjugated gear pair. Here, the generating gear is a planar gear with a pitch angle small than 90 deg, which gives the produced gear a little relief in the profile direction. In the standard planar-generating gear shown in Fig. 2, the right and left blades move along lines ocP and ocQ, respectively, which are located on the machine plane and past the apex point. The loci of the blades form the teeth of the generating gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000083_ssd.2019.8893176-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000083_ssd.2019.8893176-Figure1-1.png",
        "caption": "Fig 1. Khepera II robot schema",
        "texts": [
            " KHEPERA II MODEL AND ITS KINEMATIC CONTROL One of the most important challenges in robotic field is the autonomous motion control. The most important requirement for mobile robot is how making him self-ruling and autonomous. Then, the robot goal is being intelligent and having a capacity to overcome any environment changing. In this work, we are based on the Khepera II mobile robot. This considered robot is supported by two independent wheels. The control of this type of robot is allowed by acting on each wheel speed. 1. Khepera II kinematic model The khepera II robot is presented in figure 1 in its environment. The kinematic model of the Khepera II robot [12], [13] is presented by: ( )cos 2 d gV V x \u03b1 + = ( )sin 2 d gV V y \u03b1 + = (1) d gV V L \u03b1 \u2212 = V d and gV are right and left velocities of the two wheels. L is the distance separating the left and right wheels of the robot. (x, y,\u03b1) are the robot localization and orientation (see figure 1). Where: \u2022 , , : is the coordinate frame So, now our objective is how to obtain right and left wheel\u2019s mobile robot velocities to reach a desired destination considering an environment without and with obstacles. The control loop of the Khepera II robot is depicted in figure 2. 2. PI Kinematic Controller In navigation task, the robot must achieve the desired target location. The kinematic Khepera II model (1) can be written as: ( ) d g x V y A V \u03b1 \u03b1      =           (2) With ( ) ( ) ( ) ( ) ( ) cos cos 2 2 sin sin 2 2 1 1 A L L \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 = \u2212                  (3) The error vector which is the difference between actual and desired final positions can be defined as: x xe dx e y yy d e d\u03b1 \u03b1\u03b1 \u2212 = \u2212 \u2212                (4) So, by adopting the kinematic model, the issue is to arrive at adequate speeds T V Vgd     essential to make the mobile robot navigate from a start point coordinates [ ]Tx y \u03b1 to a desired one with coordinates T x y d d d \u03b1    "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001515_acc.2014.6859324-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001515_acc.2014.6859324-Figure3-1.png",
        "caption": "Fig. 3: Tilt and twist error definitions for the resolved tilt twist method",
        "texts": [
            " [7] to compute the attitude error in a manner that suited tailsitter dynamics better than the QF method. A slightly modified version of the RTT method is presented here. Unlike the QF method which treats the attitude error as a single rotation, the RTT method divides the attitude error into two components: tilt, which depends on the error about the body frames y and x axes, and twist, which accounts for the heading error. The tilt error, \u0398tilt, is calculated first, and is the angle between the desired and current x axis as shown in Figure 3a. Matsumoto et al. compute \u0398tilt by looking at the error about the current body y and z axes separately. The error about the y axis can be computed by \u0398y = atan2 (d31, d11) , where d11 = i iTb i id and d31 = k iTb i id are the projections of the current x and z axis onto the desired x axis respectively. Likewise, the error about the z axis can 2 be computed by \u0398Z = atan2 (d21, d11) , where d21 = j iTb i id. Note that d11, d21, and d31 are elements of the direction cosine matrix that describes the rotation between the current and the desired attitude given by D = KiT d Ki b =  d11 d12 d13 d21 d22 d23 d31 d32 d33  . The total tilt error is \u0398tilt = \u221a \u03982 Y + \u03982 Z . The twist error, as shown in Figure 3b, is defined to be the error about the x axis once \u0398tilt has been compensated for. To compensate for \u0398tilt, the current body x axis needs to be aligned with the desired x axis without rotating about the current body x axis. This rotation is computed using Rodrigues\u2019 rotation formula Rv = I3 + v\u0302 sin(\u0398tilt) + v\u03022 (1\u2212 cos(\u0398tilt)) , (3) where I3 is the 3x3 identity matrix, v = [vx vy vz] T is the unit vector describing the axis of rotation computed by the normalized cross product v = i ib \u00d7 i id\u2225\u2225i ib \u00d7 i id \u2225\u2225 (4) and v\u0302 =  0 \u2212vz vy vz 0 \u2212vx \u2212vy vx 0  "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001627_chicc.2014.6896037-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001627_chicc.2014.6896037-Figure5-1.png",
        "caption": "Fig. 5: Static stability/instability conditions",
        "texts": [
            " The total drag coefficient, in case of range correction (body + opened drag ring), is also shown in Fig. 3. Any spinning object will have gyroscopic properties. In spin stabilized projectile, the center of pressure is located in front of the center of gravity. Hence, as the projectile leaves the muzzle it experiences an overturning movement caused by air force acting about the center of mass. It must be kept in mind that the forces are attempting to raise the projectile\u2019s axis of rotation. Gyroscopic stability ensures that a projectile will not immediately tumble upon leaving the muzzle [2]. In Fig. 5, two cases of static stability are demonstrated: in the top figure, CP lies behind the CG so that a clockwise (restoring) moment is produced. This case tends to reduce the yaw angle and return the body to its trajectory, therefore statically stable. Conversely, the lower figure, with CP ahead of CG, produces an anti-clockwise (overturning) moment, which increases a further and is therefore statically unstable. It also possible to have a neutral case in which CP and CG are coincident whereby no moment is produced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003833_b978-0-12-800054-0.00026-5-Figure26.5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003833_b978-0-12-800054-0.00026-5-Figure26.5-1.png",
        "caption": "Figure 26.5 Example of complex shape made by laser powder bed AM.",
        "texts": [
            " The microstructure and properties of material deposited by AM powder\u2013based methods is not well characterized at the present time, but since it also starts with powder, it is logical to suggest the microstructures will be similar. \u2022 Feedstock: The powder-based AM processes almost all use spherical PA as the raw material source, with one known exception [24]. Powder quality and cost are of central importance for these processes also. \u2022 Shape Making: The shape-making capability of the powder-based AM methods is nothing short of spectacular, as can be seen from Figure 26.5. \u2022 Capital Cost: The contribution of the cost of capital to each component made via AM is significant, but is largely offset by the cost saving gained because no tooling is required. In principle, these methods also have the capability of much faster turnaround time (TAT) because the lead time for a forging die can be as long as 6 months. The laser and EB melting machines cost at least $500K. As in most cases, the actual cost depends on the size and configuration of the machine. At present, essentially all of the laser and EB machines have been installed in R&D facilities and would be overconfigured for a focused production activity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000934_1.4024781-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000934_1.4024781-Figure4-1.png",
        "caption": "Fig. 4 Shapes and dimensions of the flange and gasket. (a) Flange. (b) Gasket.",
        "texts": [
            " On the other hand, at a contact pressure of 62 MPa, the noncontact regions disappeared as shown in Fig. 3(b), and the circumferential leakage flow will be predominant. The thin polymer film method is very useful for determination of the critical contact pressure. However, the polymer film sometimes broke, as shown in Fig. 3(b) because of severe penetration by the hard ridges on the flange surfaces. Thus, the other observation method is needed for determination of the critical contact pressure. 3.1 Specimens and Experimental Apparatus. Figure 4 shows the shapes and dimensions of the carbon steel flange and copper gasket. The steel flange was of the raised face type with a flat mating face. The copper gasket was of the ring type. The dimensions of the specimens were determined almost in accordance with the Japanese industrial standards, JIS B2220 [17], with a nominal gas pressure of 0.98 MPa. The Vickers hardness of the flange and gasket were 2290 MPa and 420 MPa, respectively. Figure 5 shows a cross-section of the experimental apparatus for the leakage test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002051_s12555-014-0455-z-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002051_s12555-014-0455-z-Figure1-1.png",
        "caption": "Fig. 1. Two inverted pendulums connected by a spring [26].",
        "texts": [
            " The V\u0307i \u2264 eT i (Ai \u2212LiCi) T P1iei + eT i P1i(Ai \u2212LiCi)ei + eT i P1iFi + eiP1iFT i +2(\u2206i(x,x(t \u2212di(t)))T GT i P1iei \u2212 \u03b5iFT i Fi + \u03b5i(\u03b32 i1eT i ei + \u03b32 i2eT i (t \u2212di(t))ei(t \u2212di(t))+ \u03b3i2\u03b3i1eT i ei(t \u2212di(t))+ \u03b3i2\u03b3i1eie T i (t \u2212di(t))) + eT i P1iEi\u03c8i\u03b8i(t)+(\u03c8i\u03b8i(t))T ET i P1iei \u2212\u00b5i(\u03c8i\u03b8i(t))T \u03c8i\u03b8i(t)+\u00b5i(\u03b8 2 i0\u03b32 i3eT i ei) + eT i P2iei \u2212 (1\u2212 \u03c4i)eT i (t \u2212di(t))P2iei(t \u2212di(t))+hie\u0307T i P3ie\u0307i \u2212 (ei(t)\u2212 ei(t \u2212di(t)))T P3i hi (ei(t)\u2212 ei(t \u2212di(t))) (39) figure of two inverted pendulums connected by a spring is shown in Fig. 1. x\u030711 = x12, x\u030712 = ( m1gr J1 \u2212 kr2 4J1 )sin(x11(t \u2212d1(t))) + kr(l \u2212b) 2J1 + u1 J1 + kr2 4J1 sin(x22), x1 = [x11,x12] T , y1 =C1x1, x\u030721 = x22, x\u030722 = ( m2gr J2 \u2212 kr2 4J2 )sin(x21(t \u2212d2(t))) + kr(l \u2212b) 2J2 + u2 J2 + kr2 4J2 sin(x11), y2 =C2x2, x2 = [x21,x22] T , (42) where xi j, i, j = 1, 2 are the states of each subsystem and yi, i = 1, 2 depict the output of each subsystem. m1 = 2 kg and m2 = 2.5 kg, J1 = 0.5 N.S m2 and J2 = 0.625 N.S m2 , k = 100 N/m, r = 0.5 m, l = 0.5 m, and g = 9.81 m/ s2m/s, and b = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure11.8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure11.8-1.png",
        "caption": "FIGURE 11.8 Further possible random steps, following Figure 11.7.",
        "texts": [
            "7, a circle has been drawn about the starting point. The radius of the circle is chosen such that it is tangent to the boundary line (electrode) closest to the starting point. A temporary square grid is established as shown by the arrows, with the starting point and the tangent point of the circle establishing the gridpoint separations. The random walker now takes a step from the starting point to one of the four gridpoints. She has a 14 probability of landing on an electrode and being extracted. Figure 11.8a shows one of the other possibilities, the next circle, and the next temporary grid. In this case the gridpoints do not lie along the x and y axes\u2013\u2013 and there is no need for them tobe along these axes. Figure 11.8b shows apossible third circle and set of gridpoints. There is always a 25% probability that the walker will be extracted, this time at the outer electrode. After three steps, the probability of having been extracted is greater than 50%: p3 = 1\u2212 0:75\u00f0 \u00de3 = 0:578 \u00f011:16\u00de This probability climbs rapidly with the number of steps taken. After six steps, it is p6 = 1\u2212 0:75\u00f0 \u00de6 = 0:955 \u00f011:17\u00de The average walk doesn\u2019t take long. Computationally, computer time needed for these shorter walks is offset by the complexity of finding the closest wall and calculating the 11"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003004_pesgre45664.2020.9070363-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003004_pesgre45664.2020.9070363-Figure6-1.png",
        "caption": "Fig. 6. Meshed FE model of a BLDC motor showing uniformly demagnetized magnetic flux density (BM) profile.",
        "texts": [
            " (16) where ea is the back-EMF of phase A or the voltage induced in the stator coil, is a stator flux linkage in phase-A winding, is the angular displacement or rotor position and \u03c9m is the mechanical speed respectively. The change in ea, and BM therefore shall make these factors ideal to diagnose uniform demagnetization faults in BLDC motor. SITF and demagnetization faults are emulated together in the motor and the adverse effect in the machine quantities are observed which distinguishes from each other when these faults were accounted individually. The meshed model of a machine under SITF and demagnetization is shown in Fig. 6. The subsequent section discusses the outcomes of the faults. IV. RESULTS AND DISCUSSIONS Motor Current Signature (IPh): The SITF fault in a motor winding, results in the corresponding change in the phase currents. The no-load current of 2.7 A suddenly rises to approximately 480 A validating (11) in a very short interval of time as can been seen from Fig. 7. (a). Such a sudden rise in current signifies the insulation breakdown between the windings. However, in case of demagnetization fault, the rise in current is not so significant as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure4-1.png",
        "caption": "Fig. 4 Cutter blade",
        "texts": [
            ", 22 deg) to the generating plane, forming a tooth of the imaginary generating gear, whose cutters then infeed to the required tooth depth and generate the work gear. This cutting process, however, has no lengthwise movement along the gear\u2019s face width. When the profile angle is such that the rotating cutting blades by the cutter axis form a cone, however, a generating gear with concave teeth is produced that is used to cut convex work gear teeth. This observation implies that lengthwise crowning can be achieved based on the design of the blade profile angle. Figure 4 illustrates the geometry of the cutter blade, which includes a straight line (r \u00f0l\u00de l ) and a circular-arc tip fillet (r \u00f0f \u00de l ). The cutter 021007-2 / Vol. 138, FEBRUARY 2016 Transactions of the ASME Downloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use parameters are the profile angle ab, the cutter radius r0, the angle of tip ac, and the fillet radius qb. Rotating the cutter blade around the z axis forms the cutter surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001189_wcica.2014.7053623-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001189_wcica.2014.7053623-Figure1-1.png",
        "caption": "Fig. 1 Structure of Beaver AUV",
        "texts": [
            " According to the above theoretical analysis, in the presence of faults, external disturbances or other uncertainties in AUV, the proposed adaptive backstepping sliding mode algorithm can accommodate the unknown thruster faults and make the trajectory error converge to zero. Simulations studies were performed in order to demonstrate the effectiveness of the proposed adaptive backstepping sliding mode control algorithm regardless of the thruster faults occur or not. The underwater vehicle in the simulations is BEAVER AUV[13] developed by the authors\u2019 laboratory, shown in Fig.1. In the simulations, only the surge velocity and heading angle are controlled. The BEAVER AUV has one pair of horizontal thrusters driving the surge and yaw motion. Transform BEAVER AUV dynamic model for the degrees of yaw and y-direction into state space equation expressed by (2), yielding ( ) ( ) 1 1 2 21 1 2 2 22 2 , , x A x x u A x B f x u d x u x u = \u23a1 \u23a4 \u23a1 \u23a4 = + + +\u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 \u23a3 \u23a6 (28) where [ ]1 1 0A = 2 0.0856 0 0 0.0939 A \u2212\u23a1 \u23a4= \u23a2 \u23a5\u2212\u23a3 \u23a6 0.0591 0.0591 0.0238 0.0238 B \u2212\u23a1 \u23a4= \u23a2 \u23a5 \u23a3 \u23a6 ( ) 21 21 22 22 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000718_ecce.2015.7310072-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000718_ecce.2015.7310072-Figure3-1.png",
        "caption": "Fig. 3. One leg of an inverter. Fig. 4. Voltage drop across a switch.",
        "texts": [
            " sin 2 dsh r s qsh r r dsh qsh s inj sh r dsh qsh L n T n L L L T v n L L L \u03b8 \u03b8 \u03b8 \u03b8 \u2212 = \u22c5 + \u2248 \u22c5 \u0394\u2212 r r dqsh dqshi R( ) v (6) From q-axis current ripple in (6), the position error signal can be computed by (7). \u02c6 , * 1[ ] [ ] 2 [ - 2] dsh qshr r est qsh r s sh inj L L n i n T L v n \u03b8 \u03b8= \u0394 \u22c5 \u2212 \u2248 \u22c5 \u0394 , (7) Then, the rotor position and speed can be estimated from the observer shown in Fig. 1. The observer adjusts r\u0302\u03b8 in the direction that q-axis current ripple which corresponds to the position estimation error is getting diminished. The overall procedure of the signal processing in (3)-(7) is described in Fig. 2. III. POSITION ERROR INDUCED BY INVERTER NONLINEARITY A. Inverter nonlinearity effects Fig. 3 shows one leg of an inverter feeding x-phase of a motor. In the figure, x indicates an arbitrary phase among a, b, and c. Note that there are naturally parasitic capacitors denoted as Cp which are connected in parallel with IGBTs. Thus, collectoremitter capacitance, Cce, is equal to Cp, i.e., Cce = Cp. There are mainly two switching states for an output terminal of an inverter: \u201chigh\u201d state where ixs flows through the upper IGBT or diode and \u201clow\u201d state where ixs flows through the lower IGBT or diode",
            " Tlow and Thigh indicate \u201clow\u201d and \u201chigh\u201d state durations in a sampling period, respectively. Fig. 5 illustrates _swxnv\u03b4 according to the current which is extracted from the datasheet of an IGBT module, product number PM50CL1A060. In the figure, the curve with the legend \u201caverage\u201d is calculated under the assumption of half duty ratio. The second inverter nonlinearity effect is the voltage distortion due to the dead time described in Fig. 6. In the figure, the current conduction path at each period is also drawn. During the dead time where both the upper and the lower IGBTs in Fig. 3 are off, ixs can\u2019t flow through an IGBT but can flow through a diode or the collector-emitter capacitors of both IGBTs. For this reason, the pole voltage, vxn(t), cannot instantly change in a desired way, which induces the pole voltage error. To be more specific, ixs flows through a diode during the dead time unless that diode is reverse biased. In this case, vxn(t) is simply delayed by the dead time, Td, which is shown at on sequence in Fig. 6(a) and off sequence in Fig. 6(b). This delay leads to volt-sec loss or gain denoted as the area Al or Ag, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003033_iet-epa.2019.1019-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003033_iet-epa.2019.1019-Figure2-1.png",
        "caption": "Fig. 2 Calculation results of rotor stress and deformation under 18,000 r/min",
        "texts": [
            " Taking safety factors into consideration, the strength condition of the Mises yield criterion is represented as \u03c3r4 = (\u03c31 \u2212 \u03c32)2 + (\u03c32 \u2212 \u03c33)2 + (\u03c33 \u2212 \u03c31)2 2 < [\u03c3] (6) where [\u03c3] is the allowed stress. The allowed stress can be calculated by [\u03c3] = \u03c3s/ns, where ns is the safety factor. The cold-rolled non-oriented silicon steel sheet, B20AT1200, with yield strength of 395 MPa, is adopted to manufacture the iron core of the rotor. The stress distribution and rotor deformation cloud diagram are calculated by FEA, as illustrated in Fig. 2, when the motor operates at the rated speed of 18,000 r/min. Consistent with the theoretical analysis mentioned above, the maximum stress locates at the inner side of the rotor which is 134.7 MPa and is around half of the yield strength of the stain steel sheet. The maximum deformation appears at the poles of the rotor, which is 0.0148 mm and much smaller than 10% of the air gap length, ensuring the safe operation of the motor. Furthermore, it has been pointed out in the related standard IEC 60,034-1 that the AC motor should be able to endure 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003042_j.matpr.2020.04.082-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003042_j.matpr.2020.04.082-Figure5-1.png",
        "caption": "Fig. 5. Flux lines plot of configuration 1 for I = 15A.",
        "texts": [
            "Six SOMALOY 1000 3P, 800 Mpa, pre-form blanks were procured three blanks when glued together formed the material for stator fabrication while the remaining three blanks when glued together formed the material for rotor fabrication. The properties of SMC material are sourced from Hoganas AB of Sweden while procuring the prefrom SMC blanks from which the B-H curves were drawn.The BH curves of M19 steel and SMCmaterials are shown in Fig. 3, which reveals that although the SMC has inferior relative permeability when compared with lamination steel it still posses the following desirable characteristics [4] (see Fig. 4, Fig. 5). 1. High power density by 3D magnetic flux conduction 2. Lower core losses at elevated frequencies in comparison with electrical steel 3. Good formability; complex shapes can be directly compacted without destroying the material structure and resulting deterioration of magnetic properties. The static and dynamic electromagnetic characteristics [1,5] have been obtained through FEA. MagNet 6.1.1 Finite Element Analysis software package has been employed for this while for thermal and vibration analysis ANSYS 10, another multi-physics FEA software package has been used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002607_yac.2016.7804934-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002607_yac.2016.7804934-Figure4-1.png",
        "caption": "Fig. 4. Picture of real products.",
        "texts": [
            " Section II discusses the principles of the twin-engine flying wing aircraft and gives the definitions of attitude angles in the two modes. Section III introduces the components of the aircraft and gives the transition logic and the target angle calculation algorithm. Section IV presents the experimental results and Section V presents the conclusions. This research was supported by Deep Exploration Technology and Experimentation Project under Grant No.201311194-04. The aircraft structure diagram is shown in Fig. 3, and the real product is shown in Fig. 4. The fixed-wing fuselage\u2019s layout is blended wing body layout, on the top of the fuselage two brushless motors are installed for providing the hover lift in rotary-wing mode and pull in fixed-wing mode, at the bottom of the fuselage two control surfaces are installed for controlling the vehicle attitude. To use the rotary-wing mode as a reference, the establishment of the body coordinate system F: OXYZ, the original point is the aircraft\u2019s center of gravity, OX axis perpendicular to the wing plane pointing to forward (perpendicular to the paper inwards), OY axis pointing to right in wing plane, OZ axis in wing plane perpendicular to the plane OXY pointing down"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002633_imece2016-65784-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002633_imece2016-65784-Figure1-1.png",
        "caption": "Fig. 1. Squeeze film damper geometry and coordinate system",
        "texts": [
            " In hydrodynamic bearings, the thickness of the lubricant is very small relative to the journal radius and the bearing axial length, consequently [15]: 1. The effect of the curvature of the film is negligible; hence a linear coordinate system is used to describe the lubricant dynamics. 2. The variation of the pressure across the film is negligible (i.e. 0P y   ). The SFD configuration in this work is an open-ended (i.e. no end seals) axially symmetric damper. The geometry of the system is represented in Figure 1. According to the thin film theory, an orthogonal Cartesian coordinate system {x,y,z} is fixed in the plane of the lubricant, where the z-axis is perpendicular to the plane of motion. Furthermore, an orthogonal Cartesian system {x\u2032,y\u2032,z\u2032} translating with angular 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90984/ on 03/19/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use velocity R\u03c9 is introduced where the x\u2032-axis is perpendicular to the line connecting the centers of the inner and outer cylinders, and the y\u2032-axis is in the direction of the minimum film thickness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000521_icuas.2015.7152412-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000521_icuas.2015.7152412-Figure1-1.png",
        "caption": "Figure 1. The small-scaled unmanned tandem ducted fan vehicle",
        "texts": [
            " Bin Xu is with the Vehicle Research Center, Beijing Institute of Technology, Beijing, 100081 China. (e-mail:xubin@bit.cn) Nan Huang is with the Vehicle Research Center, Beijing Institute of Technology, Beijing, 100081 China. (e-mail:hn24365@gmail.com). has been devised by Beijing Institute of Technology (BIT). Subsequent tests on the vehicle show poor performance of rolling and yawing controllability. In order to ameliorate the ill control conditions, new ducted fan vehicles with improved control structure has been constructed. Figure 1 shows the small-scaled experimental model. The flight platform is considered to consist of two main ducted fans, two bilateral EDFs, tilting system and control system. The two main rotors are used to provide direct thrust to support the vehicle. The difference of main rotor speed is variable to exert control over pSitching channel. Similarly, the variable speed differential of both sides EDFs provides the control over rolling channel. The tilting system is devised to change the directions of the EDFs\u2019 thrust vectors conversely which offer the control over yawing channel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001562_978-3-642-37387-9_4-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001562_978-3-642-37387-9_4-Figure2-1.png",
        "caption": "Fig. 2 Six degrees of freedom of an AUV",
        "texts": [
            " u, v and w represent the forward, lateral and vertical speeds along x, y and z axes respectively. Similarly, the hydrodynamic moments on AUV will be denoted by L, M and N acting around x, y and z axis respectively. The angular rates will be denoted by p, q and r components. Dynamics of AUVs, including hydrodynamic parameters uncertainties, are highly nonlinear, coupled and time varying. According to [8], the six DOF nonlinear equations of motion of the vehicle are defined with respect to two coordinate systems as shown in Fig. 2. The equations of motion for underwater vehicle can be written as follows [9]: M\u20acq\u00fe C _q\u00f0 \u00de _q\u00fe D _q\u00f0 \u00de _q\u00fe G q\u00f0 \u00de \u00bc s \u00f01\u00de where M is a 6 9 6 inertia matrix as a sum of the rigid body inertia matrix, MR and the hydrodynamic virtual inertia (added mass) MA. C\u00f0 _q\u00de is a 6 9 6 Coriolis and centripetal matrix including rigid body terms CRB\u00f0 _q\u00de and terms CA\u00f0 _q\u00de due to added mass D\u00f0 _q\u00de is a 6 9 6 damping matrix including terms due to drag forces. G(q) is a 6 9 1 vector containing the restoring terms formed by the vehicle\u2019s buoyancy and gravitational terms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003256_1.c035739-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003256_1.c035739-Figure15-1.png",
        "caption": "Fig. 15 Surface contact model of two bodies.",
        "texts": [
            " Besides, a damping term associated with restitution coefficient is introduced to simulate the energy loss during contact using damping force. Accordingly, the normal contact force is expressed as FN K\u03b43\u22152 1 3 1 \u2212 c2e 4 \u03b4 0 \u03b4 0 \u2212 (1) The contact stiffness coefficient of impact body K is correlated with the material properties and structural shapes of contact planes, which can be expressed as K 4 3\u03c0 hi hj RiRj Ri Rj 1\u22152 (2) The equivalent spherical radii Ri; Rj for the impact between two bodies can be indicated by Fig. 15. The parameters hi; hj are related to the material properties, which can be expressed as Actually, the Hertz contact force ignores the friction on contact planes, which fails to fully express the tangential contact friction characteristics of contact planes. In this paper, the Threlfall model [18] (Fig. 16) is employed to simulate the tangential friction of cylindrical pair impact contact with clearance. Its expression is as follows: fT \u2212\u03bcfFN v\u03c4 v\u03c4 1 \u2212 e\u22123jv\u03c4 j\u2215vr (4) With the aid of the aforementioned established multiflexible bodymodel of landing-gear retraction dynamics, simulation analysis is performed separately concerning the effects of hinge clearance and axis deviation on the 3-D retraction mechanism, and model verification is carried out based on the test results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000021_chicc.2019.8865754-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000021_chicc.2019.8865754-Figure1-1.png",
        "caption": "Fig. 1: The coordinate reference frames of quadrotors",
        "texts": [
            " In section 2, the dynamic model for the quadrotor is set up. Subsequently, controller development is detailed in section 3. The simulation results and the analysis have been shown in section 4. Finally, short conclusions are presented in section 5. The traditional structures of the quadrotor can be divided into \u201d\u00d7\u201d mode and \u201d+\u201d mode. We choose the \u201d+\u201d mode to establish the model of quadrotor. In order to establish the quadrotor system dynamic model, the Earth-fixed inertial frame N (OnXnYnZn) and the Body-fixed frame B (obxbybzb) are defined as Fig. 1 does. The Earth-fixed inertial frame is fixed on the ground with the quadrotor take-off point as the origin On, the axis Xn points to the geographical north, the axis Yn points to the geographical east and the axis Zn points to the ground. The Body-fixed frame is fixed on the quadrotor body. Without losing the generality of the dynamic model of the quadrotors, we make the following assumption. \u2022 The quadrotor is supposed to be rigid with a completely symmetrical structure. The elastic deformation won\u2019t occur during the flight"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001088_s11044-014-9417-8-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001088_s11044-014-9417-8-Figure15-1.png",
        "caption": "Fig. 15 (Left) A humanoid robot transitions from unconstrained task-level motion control of its hands to constrained task-level motion and force control of an object that has been grappled. Given a model of the environment interaction constraints would be inferred. Constraint maintenance conditions would then be specified as inequality conditions (e.g., minimum normal forces needed to be applied by the hands on the object to maintain grapple) and equality conditions on the constraint forces would be specified to the controller. (Right) A humanoid robot generates a desired force at the hand (against a wall, for example) while maintaining desirable contact forces at the feet",
        "texts": [
            " A set of equality conditions on the robot/environment constraint forces are specified that satisfy the inequality conditions associated with maintaining desired robot/environment interactions (e.g., holding onto an object, releasing an object, maintaining contact with surfaces). These conditions are aggregated with any control conditions on the internal mechanism constraints of the robot and sent to the controller. The resulting conditions on the constraint forces, along with the specified task and null space motion commands are used to compute the control input (joint torques) to the robot. Figure 15(left) depicts a simulation scenario where the humanoid robot transitions from unconstrained task-level motion control of its hands to constrained task-level motion and force control of an object that has been grappled. The methodology of Fig. 14 can be applied to this problem. Given a model of the environment the constraints associated with interacting with the grappled object would be inferred. Constraint maintenance conditions would then be specified as inequality conditions (e.g., minimum normal forces needed to be applied by the hands on the object to maintain grapple). Equality conditions on the constraint forces would then be specified to the controller, consistent with the constraint maintenance conditions. A similar methodology could be applied to a humanoid robot interacting with the environment through foot/ground contact and hand contact. Figure 15(right) depicts a simulation scenario where the humanoid robot generates a desired force at the hand (against a wall, for example) while maintaining desirable contact forces at the feet. 3.4 Discussion of particular limitations on control We have to this point not discussed specific situations under which the general formulation of the control problems presented in this paper have no solutions. Specific problems can be physically ill-posed in which case rank deficiencies will exist in the system matrix",
            " As a practical matter it is assumed that the controller has access to the system state (via a forward dynamics solver in the simulated case or via sensors in the physical case) and estimates of the dynamic properties of the physical system. It should also be noted that the system of equations described by (41), (60), (71), and (75) may not always be well conditioned. That is, the system matrix in these equations may not always be invertible. Even given a satisfaction of the conditions relating the number of actuators, motion coordinates, and constraint coordinates; constraint forces may not always be arbitrarily specified. For example, given a bimanual robot holding an object (see Fig. 15), the vertical constraint forces applied to the object would need to balance the weight of the object and could not be arbitrarily specified under static equilibrium. Additionally, the normal constraint forces applied by the robot hands against the object would need to balance under static equilibrium as discussed in the previous section. Acknowledgements This work was performed under the DARPA SyNAPSE contract HR0011-09-C-0001. This article is approved for public release, distribution unlimited"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000665_1.a33330-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000665_1.a33330-Figure2-1.png",
        "caption": "Fig. 2 Inside view of the ACHILES projectile prototype. The IMU, which is on the other side of the main electronics boar, is not pictured here.",
        "texts": [
            " III, an uncertain model of the body dynamics is built using estimated models, and the actuator model is presented. Section IV details the design of the pitch autopilot, using three approaches with decreasing controller complexity. The robustness of the designed control laws is assessed in Sec. V, and in Sec. VI the proposed controllers are implemented and compared on the projectile prototype. The ACHILES projectile prototype is based on a fin-stabilized 80-mm-caliber shell with four nose-located canards, of which an inside view is given in Fig. 2. This missile-like structure has several advantages for identification and control investigations: as the frame is stabilized with tail fins, the stability only depends on the geometry of the aerodynamic surfaces, as opposed to a gyro-stabilized projectile. Moreover, the cross-axis coupling and the mechanical complexity are greatly reduced because there is no need for a high spin rate. The projectile is installed in the test section of a subsonic wind tunnel by the means of a 3-DoF gimbaled structure as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.28-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.28-1.png",
        "caption": "Fig. 18.28 Fiber optic gas sensor",
        "texts": [
            " Spectroscopic systems for measuring optical absorption are useful for the UV and IR wavelengths and can be used to detect many chemicals, by producing a more complex absorbance signature in the form of a spectrum. Bench-top IR instruments typically use dispersive IR techniques. In these instruments a grate or prism is used to provide a broad wavelength range to select a specific wavelength of light to pass through the sample. In all strategies, the wavelength of the light source is routinely matched to the reactive energy of the optrode indicator to achieve a best possible electronic signal. Fiber-optical chemical sensors (Fig. 18.28) use a chemical reagent or sorbent phase to alter the amount or wavelength of light reflected by, absorbed by, or transmitted through a fiber wave-guide. A fiber optic sensor typically contains three parts, a source of incident (pilot) light, an optrode, and a transducer (detector) to convert the changing photonic signal to an electrical signal. It is the optrode that contains the reagent phase membrane or indicator whose optical properties are affected by the analyte [75]. The location of the reagent, and the specific optical characteristic that is affected by it, vary from one type of the optical sensor to another"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003195_s42235-020-0063-y-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003195_s42235-020-0063-y-Figure3-1.png",
        "caption": "Fig. 3 Overall structure of the bionic flapping-wing flying robot.",
        "texts": [
            " The dimensionless form of N\u2013S, the equation in the inertial system XOYZ is as: 0, u u u X Y Z          (5) Journal of Bionic Engineering (2020) Vol.17 No. 4 where 2 2 2 2 2 2 1 , u u u u u v w X Y Z p u u u X Re X Y Z                            2 2 2 2 2 2 1 , v v v v u v w X Y Z p v v v Y Re X Y Z                            2 2 2 2 2 2 1 , w w w w u v w X Y Z p w w w Z Re X Y Z                            u, v and w are three components of the dimensionless velocity, p is dimensionless pressure and Re is Reynolds number. The overall structure of the bionic flapping-wing flying robot is shown in Fig. 3. 3 Active Disturbance Rejection Control (ADRC) control method based on expanded state observer for a bionic flapping-wing vehicle The attitude control of the bionic flapping-wing flying robot is adjusted by the deflection of the tail. The tail-wing controller controls the tail-wing attitude by two motors and adjusts the flight direction. The signal receiver is used to receive the speed and attitude control instructions given by the remote controller. Fig. 4 Principle of attitude control for a bionic flapping-wing flying robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001878_tmag.2015.2489701-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001878_tmag.2015.2489701-Figure8-1.png",
        "caption": "Fig. 8. Arbitrary-shaped coil wrapped by OGWM with six layers and four turns of each layer.",
        "texts": [
            " 7 have shown that the PEC model with the optimal \u03b5equi calculated by (K 1 + M1)/2 has the best accuracy for the calculation of the equivalent capacitance of the cylindrical coil wrapped by using both winding methods. Compared with it, the analytical formula has a percentage error more than \u221225% for OGWM and \u22126% for OCWM, respectively. Moreover, the analytical formula for OCWM can be used if the cylindrical coil is long. It is also in agreement with the assumption of well-known analytical formula. To show the efficiency and the accuracy of the presented approach, two arbitrary-shaped coils, one of which is wrapped by OCWM with four layers and the other is wrapped by OGWM with six layers (Fig. 8), are simulated. Since the simulations of both coils by using the original model are timeconsuming, the geometrical parameter settings are adapted to the acceptable computational effort. The PEC model with the obtained optimal \u03b5equi equal to 3.3531 for OGWM and 3.8477 for OCWM is used. Then, the percentage error \u03b6 of the analytical formula, the percentage error \u03be of the presented model, and the mesh time \u03c4 and the computing time t of two models related to n are evaluated in Table I. The symbols with subscripts 1 and 2 are for OGWM and OCWM, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003307_j.matpr.2020.06.269-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003307_j.matpr.2020.06.269-Figure13-1.png",
        "caption": "Fig. 13. Final Component.",
        "texts": [],
        "surrounding_texts": [
            "Zero Degree Draft Benefits: Reduced machining time, Thinner and consistent wall thicknesses, Reduced overall part weight. Unlimited Pattern Life: Guaranteed for the life of the project, Quick and easy tool modifications. Typical Size Range Tolerances Surface Finish Min. Draft Required Min. Section Thickness Typical Order Quantities Typical Tooling Costs V-PROCESS Castings - Extremely fine sand is \u2019vacuum packed\u2019 around pattern halves. The pattern is removed and metal poured into cavity. The vacuum is released and the casting removed. Up to 150 lbs 0.01000 for the first 100 , then add 0.00200 per inch. Add a maximum 0.02000 across parting line 125\u2013150 RMS None 0.12500 All $3,000 to $14,000 Sand Castings - Treated sand is molded around a wood or metal pattern. The mold halves are opened and the pattern removed. Metal is poured into the cavity. The mold is broken and the casting removed. Ounces to tons 1/3200 to 600 , then add 0.00300 per inch. Add 0.02000 to 0.09000 across parting line 200\u2013550 RMS 1 to 5 degrees 0.2500 All $800 to $4,000 Investment (Lost Wax) Castings - A metal mold makes wax replicas. These are joined and surrounded by an investment material. Wax is melted out and metal is poured into the cavity. The molds are broken and the casting removed. Ounces to 20 lbs 0.00300 to 1/400 0.00400 to 1/200 0.00500 to 300 , then add 0.00300 per inch 63\u2013125 RMS None 0.06000 Under 1000 $3,000 to $20,000 Die Castings - Steel dies, sometimes water cooled, are filled with molten aluminum. The metal solidifies, the die is opened and the casting ejected. Ounces to 15 lbs 0.00200 per inch. Add 0.01500 across parting line 32\u201363 RMS 1 to 3 degrees 0.03000 to 0.06000 2500+ $10,000 to $100,000 Ple pro Tight Tolerances: \u00b10.01000(0.25 mm) for the first 600(152.4 mm), add 0.00200(0.05 mm) per inch thereafter, add \u00b10.02000(0.508 mm) across parting line. Surface finish 125\u2013150 rms Thin Wall Sections: Standard wall thickness of 0.12500(3.175 mm), Thinner wall sections are achievable in isolated areas. ase cite this article as: J. Senthil, M. Prabhahar, C. Thiagarajan et al., Studies cess for new age castings, Materials Today: Proceedings, https://doi.org/10 4.3. V-process applications The only difference between prototype and production tooling is the additional impressions that may be added to the pattern to fully utilize the entire mold cavity [8]. Typical applications are: Medical Devices Computer on performance and process improvement of implementing novel vacuum .1016/j.matpr.2020.06.269 Instrumentation Electronic Enclosures"
        ]
    },
    {
        "image_filename": "designv11_30_0003470_j.jnnfm.2020.104406-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003470_j.jnnfm.2020.104406-Figure2-1.png",
        "caption": "Fig. 2. Schematic drawing of the circular cylinder test section and the two pressure taps used for pressure drop measurements.",
        "texts": [
            " To be more specific, about 7 mm away from the start of the 10 mm long, straight channel\u2014a distance much longer than the conventional minimum entrance length (for turbulent flows) of ten times the width of the channel (a rule of thumb in fluid mechanics) [120]. Enough length was also left downstream of the cylinders for the flow to redevelop, ensuring no exit effects either. In order to measure the pressure drop \ud835\udee5\ud835\udc43 caused by the presence of each cylinder in the flow, a couple of pressure taps were attached to the microchannel. These are located \ud835\udc3fp = 0.5 mm away and on opposite sides of each cylinder, one upstream and the other one downstream. Fig. 2 shows the position of the pressure taps. The design of the pressure taps was determined in a previous study where several configurations were tested in a plain microchannel (exactly like the one used here) to assess which one was the most accurate [114]. The intake width \ud835\udc64 of the taps is \ud835\udc64\u2217 = 108 \u03bcm and their roundish pressure sensing areas are about 1.5 mm in diameter. Fig. 3 was included to showcase the accuracy and minimal flow disturbance of the pressure taps. Fig. 3[a] gathers results of pressure drop measurements over a length of about 1 mm (the distance between the taps) in an obstacle-free replica of our microchannel using two Newtonian fluids: deionised water (DIW) and an aqueous solution of 68 wt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.33-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.33-1.png",
        "caption": "Figure 3.33 Stepper motor cross-section; hybrid",
        "texts": [
            " 82 Electromechanical Motion Systems: Design and Simulation To rotate 15 \u25e6CW, poles A and A\u2032 must be turned off and pole B must become a north pole, pole B\u2032 a south pole. The pole polarities and step sequence for a full step 360 \u25e6CW rotation is: A S000N000S000N000S000N000S A\u2032 N000S000N000S000N000S000N B 0N000S000N000S000N000S000 B\u2032 0S000N000S000N000S000N000 C 00S000N000S000N000S000N00 C\u2032 00N000S000N000S000N000S00 D 000N000S000N000S000N000S0 D\u2032 000S000N000S000N000S000N0 Half step and two phases on full step can be achieved in the same manner as for the VR motor. The hybrid stepping motor is essentially a combination of the VR and PM motors. Figure 3.33 shows both a motor cross-section and side view of the rotor. System Components 83 The cylindrical rotor consists of three main parts: A center section that is an axially magnetized permanent magnet plus two end caps that have small salient poles, or teeth, around their circumference. This construction results in one end cap becoming a south pole and the other a north pole. The teeth on one end cap are offset from the teeth on the opposite end cap by one half a tooth pitch. The stator has wound salient poles with teeth on the same pitch as the rotor teeth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000026_memsys.2019.8870765-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000026_memsys.2019.8870765-Figure6-1.png",
        "caption": "Figure 6: Prototype swimmer.",
        "texts": [
            " Compared with the U-IDT actuator in air, the admittance characteristics of the U-IDT is changed in water. The conductance of the U-IDT actuator was still around 17.5 mS. The susceptance was up to 16.5 mS. However, the 9.61 MHz driving frequency of the lithium niobate U-IDT actuator was remained in water. Two lithium niobate U-IDT actuators were fabricated by the MEMS process. To generate the acoustic jet with radiational longitudinal wave, A cavity was designed. The novel prototype swimmer (45x14x16 mm) was fabricated, as displayed in Figure 6. Two lithium niobate U-IDT actuators were set to excite the LSAW, propagating along the boundary surface between the lithium niobate substrate and water. Simultaneously, the radiational longitudinal wave was generated into the water with 22 \u030a Rayleigh angle. The cavity model was designed to make a acoustic jet with the radiational longitudinal wave in the rear part of the cavity. The novel prototype swimmer was fabricated by a 3D printer. A foam kickboard was set to make the balance in water. To evaluate the performance of the novel swimmer, the no load speed was measured in the tank filled with water"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003522_s00542-020-05045-8-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003522_s00542-020-05045-8-Figure12-1.png",
        "caption": "Fig. 12 Mode shapes under pre-stressed condition",
        "texts": [
            " The modal analysis uses only the stiffness matrix obtained from static structural analysis. Table 7 shows the modal modes under un-pre-stressed and pre-stressed conditions. The results show insignificant difference between the modal frequency values between un-pre-stress and prestressed conditions. Therefore, pre-stress effects are found to be negligible. The calculated ITS for first modal frequency is 0.000324 s. This means that an impulse force of this duration magnitude causes first resonance vibration. The mode shape for first and fourth modals are shown in Fig. 12. On the other hand, the natural frequencies of the gripper grasping an object will be significantly affected by the weight and size of the object. It is expected that the natural frequencies values will drop depending on the weight of the object. The dynamic time response of the gripper\u2019s opening and closing are simulated under EM and PE actuations. The dynamic damping properties of the ABS are assigned from the material data base in ANSYS Workbench. This section studies the deflection time response for three cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure5.14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure5.14-1.png",
        "caption": "Figure 5.14 Wafer spinner assembly components",
        "texts": [
            " In selecting a motor for the wafer spinning application, the object is to choose one that has the lowest dissipation in order to minimize the temperature rise occurring at the surface of the wafer during the coating process. Specifications: Maximum speed: 10 000 rpm = 1047 rad s\u22121 Maximum acceleration: 50 000 rpm s\u22121 = 5236 rad s\u22122 Nominal profile: Accelerate to 5000 rpm, run for 5 s, and decelerate to zero velocity. Components: A hollow shaft connecting the motor rotor to a Delrin chuck on which the silicon wafer is held in place by vacuum introduced through the shaft (see Figure 5.14). System Examples \u2013 Design and Simulation 273 Inertias: Chuck: J = 0.359 g cm s2 Wafer: J = 0.792 g cm s2 Shaft: J = 0.180 g cm s2 Total load inertia: Jl = 1.33 g cm s2 Acc/Dec times: 0.1 s to 5000 rpm; 0.2 s to 10 000 rpm At the high speeds and acceleration involved, a series of motors must be evaluated with respect to dissipation caused by I2R during acceleration and deceleration and core loss dissipation occurring while running at 5000 and 10 000 rpm. Based on bearing andmounting requirements, a size 90mmmotorwas chosen for evaluation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003623_ecce44975.2020.9235618-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003623_ecce44975.2020.9235618-Figure6-1.png",
        "caption": "Fig. 6. Mechanical model of 18-12 pole SRM",
        "texts": [
            "10 A TABLE II PARAMETER OF PROPOSED CURRENT AT 1000 RPM, 10 NM Parameter Proposed Current i0 (A) 11.4 i (A) -12.5 i (A) 1.47 i (A) -1.765 (deg) -20 (deg) 300 (deg) 240 Fig. 4. Radial force sum waveforms of the square and the proposed current. 0 200 400 600 800 0 0.001 0.002 0.003 0.004 0.005 R ad ia l F or ce S um ( N ) Time (s) FEM - Square Current FEM - Proposed Current 616 N158 N 10 Nm 1000 rpm FEM analysis 4691 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 15:23:01 UTC from IEEE Xplore. Restrictions apply. Fig. 6 shows the mechanical model of the 18-12 pole SRM. The model is used for the mechanical analysis in ANSYS Mechanical. The 18-12 pole SRM parts consist of front cover, the housing, the rear cover, three bolts, coils, the rotor, bearings, the shaft, and the stator. The bolts are used to fix the stator and the housing. On the housing, there is a hole for the vibration pick-up sensor. The outer surface of the front cover is mounted to the experimental test bench. The front cover is mounted by using six bolts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002473_j.measurement.2016.11.034-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002473_j.measurement.2016.11.034-Figure9-1.png",
        "caption": "Fig. 9. Impact test configuration for the compensation approach validation.",
        "texts": [
            " For the points not meeting this criteria, the coupling between the 2 axis involved was set to zero. As showed by these results the dynamic compensation allows a device bandwidth extension up to 10 kHz. The compensated cross-sensitivities are also reduced to a negligible level for a large part of the studied frequency band. Impact tests were next performed to study the influence of the correction method for a known short-lived excitation. During these tests, the uniaxial force sensor mounted on the calibration tool was hit in each of the three directions of measurement by a steel ball (Fig. 9). This configuration allowed short-duration impacts down to 150 ls. At the same time, the input force applied and the 3 output components of force were recorded. The same accelerometric compensation as in Section 3 was applied to the input force signal. These validation tests were carried out individually for the 3 directions of the experimental device to validate correction of each measurement axis. Fig. 10 shows the reference input force (Fref ) and the raw (Fraw) and corrected output forces (Fcorr)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002629_smc.2016.7844333-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002629_smc.2016.7844333-Figure6-1.png",
        "caption": "Fig. 6 2DOF PID pitch attitude control system",
        "texts": [
            " CONTROLLER DESIGN In order to avoid actuator saturation, one design procedure is to calculate the proportional gain considering the actuators limit and the maximum anticipated input error and use the derivative gain to improve the transient response of the system and finally apply integral control to improve the steady state error if there is any residual error due to disturbance, approximation or linearization [1]. However, this method doesn\u2019t optimize the input. Here, instead we will use the 2DOF PID controller architecture shown in Fig. 6. 1) Proportional and Integral Parameters: In any control a smooth transition to the reference set point is desirable. One way of minimizing overshooting the set point when using a PID controller [9] is to let the proportional parameter act of the output feedback while letting the integral parameter act on the error feedback. This way a less violent controller for a step change in the input can be designed. 2) Derivative Parameter: For aerodynamic stability, a derivative control is needed to damp pitch angle oscillation, but applying derivative control on the output amplifies error",
            " Two delayed reference inputs ( \u2212 1), ( \u2212 2), one delayed controller input ( \u22121) and two delayed outputs ( \u2212 1), ( \u2212 2) with a sampling period of 0.05 are used as inputs to NN controller. Three layers with 5 neurons in the input, 13 neurons in the hidden layer and 1 neuron in the output layer are used as NN architecture. 2000 samples with input/output maximum/minimum size of 0.7854/\u22120.7854 radians respectively are used for controller training. In both neural nets, the LevenbergMarquadt algorithm is used for learning and parameter adjustment. The NN and the 2DOF PID controllers were simulated in Simulink as shown in Fig.7 and Fig.6 respectively. The controllers set-point tracking ability were tested using a step change in an input signal of = 15 degrees. Fig.2 shows the PID controller tracks the set point smoothly with a minimal overshoot less than 0.02 degree. The system response has a settling time less than 4.5 seconds. Concerning the actuator effort the designed PID controller performs well even when used with the full longitudinal system which requires less than 5-degrees elevator deflection as compared to the approximated model which requires 11-degrees elevator deflection to achieve the desired pitch attitude as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002926_s1068798x20010141-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002926_s1068798x20010141-Figure1-1.png",
        "caption": "Fig. 1. Analysis of measurement errors in the calibration of a force sensor (a); errors in measuring the introduction area of a spherical waveguide tip (b); and errors in probe measurements (c).",
        "texts": [
            " In any case, the axis of the probe containing the optical waveguide will unavoidably deviate from the normal to the soft-tissue plane by an angle \u03b1. Therefore, the error in pressure measurement associated with this deviation must be analyzed. In manual probe supply to ensure contact with the soft tissue, \u03b1 will be considerably greater than in a robot system. That must be taken into account in selecting the numerical values of \u03b1. The angular deviation is systematic and may be broken down into several components: a) the measurement error \u03b4k associated with calibration of the force sensor when the force does not act along the normal (Fig. 1a); b) the error \u03b4D in measurements using the probe when the force does not act along the normal (Fig. 1c); c) the error \u03b4S in the area of waveguidetip introduction in the soft tissue (Fig. 1b). The irregularities of the soft tissue are ignored, since the pressure is measured with considerable depth of the waveguide tip in the soft tissue (more than 0.050 mm). In such conditions, the soft tissue wraps itself around the spherical waveguide tip. Analysis of the errors is illustrated in Fig. 1. ( ) ( )P P F F S S\u00b1 \u03b4 = \u00b1 \u03b4 \u00b1 \u03b4 RUSSIAN ENGINEERING RESEARCH Vol. 40 No. 1 In the present case, the error \u03b4F in measuring the force consists of several components: (1) the estimated error associated with calibration of the force sensor; (2) the estimated error in measurements using the probe. The first component is expressed in the form \u03b4k = (1 \u2013 cos\u03b1) \u00d7 100% Assuming that \u03b1 is approximately 1\u00b0, we obtain \u03b4k = 0.015%. The second error takes the form \u03b4D = (1 \u2013 cos\u03b1) \u00d7 100%. Assuming that \u03b1 is approximately10\u00b0, we obtain \u03b4D = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001233_amr.874.57-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001233_amr.874.57-Figure1-1.png",
        "caption": "Fig. 1. Overview (a) and 3D model (b) of the prototype hydraulic TPM: 1 \u2013 fixed base, 2,3,4 \u2013 rotary mounts of hydraulic cylinders, 5,6,7 \u2013 integrated electro-hydraulic axes, 8.9.10 \u2013 joints of moving platform, 11 \u2013 pin joints.",
        "texts": [
            " Parallel manipulator are computer-controlled mechatronic devices, which have been used to assist humans in various tasks [1,2]. In Department of Mechatronic Devices at Faculty of Mechatronics and Machine Devices of Kielce University of Technology (Poland) was constructed the three axis (3-axis) and three-degrees-of-freedom (3-DoF) hydraulic translational parallel manipulator (TPM), that consists of a fixed base, moving platform and three integrated electrohydraulic axes, connected by the joints. Overview and 3D model of the prototype hydraulic TPM was shown in Figure 1. Three equal integrated electro-hydraulic axis create prismatic joints P of the hydraulic TPM [3]. A single electro-hydraulic axis consists of CS type cylinder internally integrated with the magnetostrictive linear position sensors (Novostrictive\u00ae) and externally integrated with 4/3-way high response directional valve directly actuated with electrical position feedback of type 4WRSE. Hydraulic cylinder parameter: piston diameter D = 40 [mm], rod diameter d = 28 [mm], stroke All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of TTP, www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001205_j.compstruc.2013.03.011-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001205_j.compstruc.2013.03.011-Figure3-1.png",
        "caption": "Fig. 3. Clamped-pinned elastic linkage loaded (a) by a horizontal force P and (b) by a follower force F at the roller. The primary structures corresponding to the problems are formed by removing the roller at hinge \u2018\u20180\u2019\u2019 and replacing it with the redundant force (a) Ry and (b) R y .",
        "texts": [
            " (7) and det (A) = 0. This procedure needs less computati onal effort than the consecutive calculations for the equilibriu m paths and then for the critical points of those paths. We demonstrate the usage of this method via computin g the critical equilibrium configurations of a clamped-pinne d, non-linea rly elastic linkage subjected to simple end loads in the next section. We investigate a linkage that is elasticall y clamped at one end and supported by a roller at the other end, as shown in Fig. 3. Two types of loads are applied: the linkage is loaded at the roller either by a force acting in the horizontal direction (Fig. 3 (a)), or by a non-potential follower force (Fig. 3 (b)). Therefore, methods based on the total potential energy function for finding equilibrium states and critical equilibriu m configurations cannot be applied. We emphasise that in case of non-potential loads the stability of equilibriu m states and different types of dynamical loss of stability can only be explored by a dynamical analysis. We suppose that the moment in the ith spring is written in the form mi ai ai 1 j0 i \u00bc qim\u0302i ai ai 1 j0 i . Here qi is the initial stiffness of the ith rotational spring, and the function m\u0302i\u00f0n\u00de is odd, is bijective, is smooth and at least once differentiab le in a given domain of n, and it has a unit tangent at n = 0: dm\u0302i\u00f0n\u00de=dnjn\u00bc0 \u00bc 1",
            " the springs are softening (or sublinear) in a special way. The redundant moment MR is zero at hinge 0, thus j0 0 \u00bc 0 and m0(a0 a 1) = 0. This, alongside with the boundary condition a 1 /right = 0, implies that a 1 = /right = a0 formally. For convenience we take j0 i \u00bc j0 for all i \u2013 0 and we set /left = 0. Besides, xN is arbitrary now. Therefore, the essential boundary condition s are: x0 \u00bc 0; y0 \u00bc 0; \u00f012\u00de yN \u00bc 0; aN \u00bc 0: \u00f013\u00de In this load case a compressive horizontal force is exerted on the linkage at the roller, as shown in Fig. 3 (a). This load case can be derived from the general one by setting H1 = P, V1 = 0, M1 = P \u2018/2 sin a0, and all the other loadings are zero: Hj = 0, Vj = 0, Mj = 0 (j = 2, . . . ,N). The statical initial values of map (5), following definitions (4), are u0 = P, p0 = 0, w0 = P \u2018/2 sin a0, q0 = 0, r0 = P\u2018/2 sin a0. Simplifying map (5) according to this special load case and introducing the dimensio nless load parameter k = P\u2018/q, the dimensionless redundant force C = Ry\u2018/q, and the dimensionle ss distances xi ",
            " Finally, we can formulate the condition for an equilibriu m state to be critical as det\u00f0A\u00de \u00bc @aN\u00f0a0;C; k\u00de @a0 @yN\u00f0a0;C; k\u00de @C @aN\u00f0a0;C; k\u00de @C @yN\u00f0a0;C; k\u00de @a0 \u00bc 0: \u00f019\u00de If we are to find the equilibrium states of the structure , we have to solve Eq. (15) for a0 and C. The results form one dimensio nal manifolds locally, the equilibrium paths. The parameter is the load k. However, if we are interested only in the critical states of the structure , then we need to solve the system of non-linear equations composed by Eqs. (15) and (19) for a0, C, k. The solution is a zero-dimens ional set, the critical points of the equilibriu m paths. In this load case a compress ive follower force acts on the linkage at the roller, as shown in Fig. 3 (b). This loading can be obtained from the general case by setting H1 = F cos a0, V1 = F sin a0, M1 = F cos a0 \u2018/2 sin a0 F sin a0 \u2018/2 cos a0 = 0 and all the other loadings are zero. From (4) we set the statical initial values u0 = H1, p0 = V1, w0 = H1 \u2018/2 sin a0, q0 = V1 \u2018/2 cos a0, r0 = M1 = 0. From (5) we can derive the map for this buckling problem. If we introduce a dimensionless load parameter k\u2044 = F\u2018/q, a dimensionle ss redundant C \u00bc R y\u2018=q, and xi ? xi/\u2018, yi ? yi/\u2018, we can write this map as xi \u00bc xi 1 \u00fe cos ai 1; yi \u00bc yi 1 \u00fe sin ai 1; ai \u00bc ai 1 \u00fe j0 \u00fe m\u0302 1 i \u00f0C xi k cos a0 yi \u00fe k sin a0 xi\u00de: \u00f020\u00de As in the previous example, map (20) can be iterated for i = 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003055_s12289-020-01560-1-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003055_s12289-020-01560-1-Figure9-1.png",
        "caption": "Fig. 9 Mean dimensional tolerances (\u00d710 \u03bcm) of all samples by experimental tests",
        "texts": [
            " One of the most important issues in the measuring dimensions of samples was to remove samples from the building platform. Samples may be deformed because of the residual stresses. Samples can be affected by the stress relaxation phenomena. Firstly, dimensional tolerances of all samples were measured at two heights. The first measurements were close to the building platform, while the second one was close to the top surfaces of samples. Both the top and bottom areas of samples were examined. Dimensional tolerances of samples at two levels are shown in Fig. 9 using the experimental measurements. Dimensional tolerances of samples at two levels are shown in Fig. 10 by FEA analysis (Autodesk Netfabb Simulation Utility). According to ISO 286, IT tolerances of samples were added. Bottom areas of samples deformed less than upper areas. Apparently, the height of samples has considerable influence on displacements. The building platform plays a fixation role, especially for the bottom areas. Bottom areas were tried to deform similarly as upper areas. However, the building platform kept those areas stationary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000767_s12206-013-0108-9-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000767_s12206-013-0108-9-Figure11-1.png",
        "caption": "Fig. 11. (a) Dynamic displacement with different Fb (Ff = 600 N); (b) velocity-displacement diagram with Fb = 4400 N (Ff = 600 N).",
        "texts": [
            " Vibration frequency of the major period was related to external loads, while vibration frequency of the minor period was related to natural frequency of the table. As Fig. 7 shows, supporting rigidity was determined by deformation of the driving worm gear pair, which caused continuously varying natural frequency of the table. When vibration amplitude of working plane increases, supporting rigidity and natural frequency decreases accordingly. When frictional force Ff is 600 N, constant current hydrostatic guideway works in unload state. Dynamic displacement and its phase-plane diagram of velocity-displacement are shown in Fig. 11. When preload Fb is 2200 N, dynamic displacement of the working plane reaches about 0.07 mm. When preload Fb is larger than 3300 N, ZT20SW can be positioned reliably and backlash is eliminated all the time, so dynamic displacement of the working plane decreases obviously. Under this circumstance, vibration of the working plane is caused by deformation of the driving worm gear pair, supporting rigidity of which reflects the most important role of dynamic characteristics of the table. The working plane vibrates around the equilibrium position, vibration amplitude of which is less than 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003484_0954406220963148-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003484_0954406220963148-Figure2-1.png",
        "caption": "Figure 2. Probability distribution function of local search flight distance before and after improvement. (a) The initial distribution function and (b) The improved distribution function.",
        "texts": [
            " The inertia weight in the later period is kept small to improve the local search accuracy. Random exploration mechanism. When the bat population uses the formula (4) for local search, the aforementioned problems will occur. In order to improve the local search ability of BA, this article introduces the Cauchy-distribution to improve the local search formula. Thus, e will subject to the Cauchy distribution, that is, e C (0, 0.5). The probability density function curve of the Cauchy distribution is shown in Figure 2. It can be seen that the value of e C (0,0.5) has two distribution characteristics. Firstly, it is mainly distributed around 0, and the probability distribution in the interval [ 1] is 70.48%, which will guarantee the algorithm search around the optimal value and improve the algorithm\u2019s deep mining ability. Secondly, there is 29.52% probability beyond [ 1] which can make the algorithm jump out of the local optimal value with a higher probability and reduce the blindness of local searching. Random step size selection strategy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001439_s102833581411010x-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001439_s102833581411010x-Figure1-1.png",
        "caption": "Fig. 1. Structure of the drive: (1) engine, (2) gear mechanism, (3) control, and m is the object.",
        "texts": [
            " At each stage of the synthesis, it is necessary to address the simplest mathematical models gradually increasing their complexity from stage to stage. It is expedient to use the normalized (dimension less) mathematical models in which the real parame ters and other characteristics of the system are trans formed to the generalized dimensionless complexes and criteria of similarity. As a result, the number of base parameters decreases, which facilitates the analy sis and synthesis of the system under consideration. The drive structure is shown in Fig. 1. The engine can be electric, hydraulic, pneumatic, or of another type. The motion is transferred from the engine to an object by the mechanism with constant or variable gear ratio. Because at the initial stage the structure of components is either not determined at all or is presented as construction analogies, it is neces sary to consider the engine first as an ideal unit and to begin the formation of the control system structure proceeding from the simplest algorithm, which is cho sen according to the drive purpose"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000564_meco.2015.7181891-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000564_meco.2015.7181891-Figure1-1.png",
        "caption": "Figure 1. The relationships between the reference frame (xr, yr, zr), the frame (xv, yv, zv) fixed in the UAV and the camera frame (x, y, z).",
        "texts": [
            " We can obtain: 1) the pole latitude p, the pole longitude p and the height of the point of wire suspension hp from the geoinformation system; 2) the UAV latitude v, the UAV longitude v and the UAV height hv from the GPS; 3) the UAV azimuth v, the UAV elevation angle v, the UAV bank angle v from the inertial navigation system. The angles describe a relationship between the reference frame (xr, yr, zr) that is North-East-Down frame and the frame (xv, yv, zv) fixed in the UAV. The origins of both frames coincide. We assume that camera is placed in the frames origin. We also assume that we know the angle c between optic axis y and yv axis. Fig.1 shows the relationships described above. The point I, which is the point projecting on image plane (Fig.2), has the coordinates ( r i r i r i zyx ,, ) in relation to the This work has been accomplished with the assistance of the Ministry of Education and Science of Russian Federation within the agreement on granting a subsidy No. 14.574.21.0012 (RFMEFI57414X0012). \u2013 159 \u2013 reference frame. The values of the coordinates can be given by: ,,360cos2 ,3602 iv r ivEiv r i Niv r i hhzRy Rx where i, i, hi respectively represent the latitude, the longitude and the height of the point I, RN is the meridian radius of curvature, RE is the transverse radius of curvature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000852_icra.2013.6631351-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000852_icra.2013.6631351-Figure1-1.png",
        "caption": "Fig. 1. Quadrotor Pitch Control",
        "texts": [
            " The paper is organized as follows: Section 1 describes the system and its modeling, Section 2 introduces the control problem and its H\u221e formulation, Section 3 gives details on the parameter estimation and Section 4 provides flight test results and related comments. A quadrotor is composed of two pairs of counter rotating propellers mounted at the extremities of a cross-shaped frame. Decreasing the rotation speed thus the thrust of a rotor while increasing the rotation speed of the opposite one permits to create a moment to control the vehicle (Figure 1). The difference of rotation speeds is denoted \u03b4\u03c9. The attitude will be represented using the Euler angles \u03c6, \u03b8, \u03c8 and their classic aeronautical denomination roll, pitch and yaw angles. A complete model of a quadrotor, including the aerodynamics of the rotors, can be found in [8]. The roll and pitch axes of a quadrotor are qualitatively identical. The main differences between those axes would usually be the moment of inertia due to the organization of 978-1-4673-5643-5/13/$31.00 \u00a92013 IEEE 5396 the embedded system and payloads"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001561_0954406214531748-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001561_0954406214531748-Figure1-1.png",
        "caption": "Figure 1. N-slice split gear pair. Figure 2. Time-varying mesh stiffness of a pair of sub-gears.",
        "texts": [
            "16 This study starts with the analysis of the total mesh stiffness of the split gears through the Fourier expansion. Then the effect of the basic parameters including the number of gear slices and contact ratio on the parametric vibration is examined based on a simple dynamic model by using the multi-scale method with an aim to obtain analytical unstable boundaries. The analytical predictions are verified by Floquet theory. The purely rotational vibrationmodel of a split gear pair under the excitation of time-varying stiffness is shown in Figure 1. In this figure, both gears consist ofN slices and each slice is spur gearwith the samewidth b/N, where b is the axial width; is the angular velocity; T1,2 are the driving and load torque; kmn (n\u00bc 1, 2, 3, . . . ,N) is the mesh stiffness of the nth slice pair; ci is the damping at the ith (i\u00bc 1, 2, 3, . . . N) gear pair, and 1,2,R1,2, and I1,2 are the rotational angles, base radius, and rotary inertias of the pinion and driven gear, respectively. The angle between adjacent sub-gears are 1\u00bc 2p/(Z1N) and 2\u00bc 2p/(Z2N), respectively",
            " The special contact ratios, however, may not be chosen in applications because of the strength requirement, the module or other restrictions, and the fluctuation of the stiffness remains, and thus the parametric instability is studied in this work. Similar to regular spur gear transmission, a split gear pair still has the problem of parametric instability. Aiming to deal with the influence of the slice number N and the contact ratio Cr on the instability, the equations of torsional motion of the model shown in Figure 1 are needed I1 1 :: \u00fecR1 R1 _ 1 R2 _ 2 \u00fe R1km R1 1 R2 2\u00f0 \u00de \u00bc T1 \u00f012a\u00de I2 2 :: \u00fecR2 R2 _ 2 R1 _ 1 \u00feR2km R2 2 R1 1\u00f0 \u00de \u00bc T2 \u00f012b\u00de Equations (12a) and (12b) can be reduced to one equation in terms of dynamic transmission q \u00bc R1 1 R2 2, i.e. m \u20acq\u00fe c _q\u00fe kmq \u00bc f \u00f013\u00de where m \u00bc I1I2=\u00f0I2R 2 1 \u00fe I1R 2 2\u00de, c \u00bc PN j\u00bc1 cj, and f \u00bc \u00f0I2R1T1 \u00fe I1R2T2\u00de=\u00f0I2R 2 1 \u00fe I1R 2 2\u00de. The instability of the system can be determined from the dynamic equation \u20acq\u00fe 2\" _q\u00fe !2 0 1\u00fe \" X1 l\u00bc1 Cleil!mt \u00fe cc \" # q \u00bc 0 \u00f014\u00de where !2 0 \u00bc km=m, \" \u00bc \u00f0kmax kmin\u00de= km \u00bc 1=\u00f02Cr \u00de, and 2\" \u00bc c=m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001204_roman.2013.6628553-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001204_roman.2013.6628553-Figure1-1.png",
        "caption": "Fig. 1. Joint configuration of the robot arm.",
        "texts": [
            " Section II explains the control transitions to resolve robot redundancy via optimization of an energy cost function subject to constraints. In addition, the strategies for dealing with physical constraints, namely kinematic singularity and self-collision avoidance, will also be described. Section III validates the effectiveness of the IK framework and finally, Section IV will present concluding remarks. II. INVERSE KINEMATICS CONTROL FRAMEWORK The robot arm that is analyzed here, consists of a spherical joint for the shoulder, a revolute joint at the elbow and a spherical joint at the wrist, as shown in Figure 1. The redundancy of the arm can be parameterized by defining the swivel angle \u03c6, the rotation angle of the elbow around a virtual axis that connects the shoulder and wrist joints [10]. In Figure 2, as the swivel angle varies with the shoulder and wrist fixed in position, the elbow traces the arc of a circle lying on a plane whose normal is parallel to the wrist-toshoulder axis. Badler [11] described the circle mathematically and the swivel angle \u03c6 by defining a normal vector of the plane as 978-1-4799-0509-6/13/$31.00 \u00a92013 IEEE 496 the unit vector n\u0302 in the direction of the shoulder to the wrist. Unit vectors u\u0302, v\u0302 and n\u0302 form an orthonormal coordinate system for the plane containing the circle, where u\u0302 and v\u0302 lie along the plane of the circle and n\u0302 normal to the plane. The unit vector u\u0302 is obtained by u\u0302 = \u2212z + (z \u00b7 n\u0302)n\u0302 \u2016\u2212z + (z \u00b7 n\u0302)n\u0302\u2016 (1) where z is the z-axis of the fixed reference frame at the shoulder in Figure 1. v\u0302 is set as v\u0302 = n\u0302\u00d7 u\u0302. The centre of the circle, c and its radius, r is given as: c = s + cos(\u03b1)L1n\u0302, r = L1 sin (\u03b1) cos(\u03b1) = L2 2 \u2212 L1 2 \u2212 \u2016w \u2212 s\u20162 \u22122L1\u2016w \u2212 s\u2016 (2) where s and w are the positions of the shoulder and wrist respectively. L1 is the length of the upper arm and L2 is the length of the forearm. n\u0302 is an unit vector that is in the direction of the shoulder to the wrist. Then the elbow position, e is given by: e = r[cos(\u03c6)u\u0302 + sin(\u03c6)v\u0302] + c (3) where \u03c6 is defined as the swivel angle that the elbow makes with the unit vector u\u0302 on the plane of the circle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000108_ffe.13138-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000108_ffe.13138-Figure2-1.png",
        "caption": "FIGURE 2 (A) Sub\u2010surface variation of stress components in the cylin minor axis of the contact, perpendicular to the cylinder axis. (B) The thr bench.[Colour figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            " If the load path does cross the limit line, Sdv is defined as the highest point on the load path, as marked in Figure 1 and expressed mathematically as follows: Sdv \u00bc max t \u03c4max m; t\u00f0 \u00de\u00bd : (5) The probabilistic models based on DVC that are described in the coming sections use Sdv to predict fatigue lives. When an elastic sphere and cylinder are pressed into contact, an elliptical contact patch is produced with its major axis along the axis of the cylinder. A typical stress distribution in the cylinder at a certain depth beneath the contact is shown in Figure 2A. Similar distributions exist at other depths as well. Figure 2A can also be interpreted as the variation of stress components experienced by a point at that depth in the cylinder due to one rolling contact passage of the sphere from left to right above it. This rationale is applied to obtain the stress history needed to implement probabilistic models based on DVC. This section briefly reviews the traditional probabilistic models used in the bearing industry and introduces the DVC\u2010based probabilistic models proposed by the authors. der for a sphere\u2010cylinder Hertzian contact",
            " To utilize these modified equations for life prediction, the values of the constants KLP, KIH, KZ, K1, K2, and K3 need to be calculated, or in other words, these equations need to be calibrated. Calibrating them requires that we have a set of experimental fatigue lives NP at hand, while also knowing the corresponding stress quantities utilized by all these probabilistic models. For instance, we need to know the maximum \u03c40 in the stressed volume for calibrating the LP model. Accelerated RCF tests were conducted using a three\u2010ball\u2010on\u2010rod test bench (refer Figure 2B) to obtain the required experimental fatigue lives NP 3. Silicon nitride (Si3N4) balls of diameter 12.7 mm and vacuum induction melt\u2013vacuum arc re\u2010melt (VIM\u2010VAR) M50 steel rods of diameter 9.5 mm were used for the test. They were run with a radial load of 1181.71 N between the ball and the rod, which generated a maximum Hertzian pressure of 5.51 GPa at the contact.10 Multiple RCF tests can be run on the same rod as shown in Figure 2B, by advancing the rod along its axis. Out of the nine nominally identical runs that were conducted, seven showed spall formations on the rod, while the remaining two were aborted due to spalling on the silicon nitride ball and lubrication shortage. An optical micrograph of a spall formed on the rod is shown in Figure 3A. In the context of life prediction under RCF, L10 life is defined as that fatigue life that is endured in at least 90% of the runs conducted in a test. Using a Weibull chart (Figure 3B), the L10 life was estimated to be 17",
            " Figure 5 shows contour plots of the contact pressure generated at the interface and the von Mises stress generated in the rod. A mesh convergence check conducted on this model has been detailed in Appendix A. acks. (B) The Weibull chart used by Londhe et al. to calculate the L10 To obtain the macroscale stress history \u03a3(x,t) that points at a certain depth in the rod experience under one stress cycle, the six components of stress were extracted along the paths of constant depths beneath the contact as shown in Figure 4. From the extracted data, stress histories as shown in Figure 2 can be constructed for each depth. Since Sdv at a point depends solely on the stress history experienced by it, those at the same depth will experience the same Sdv. We evaluated Sdv only for the various depths in the cross section at the centre of the elliptical contact. Subsequently, when the integral is evaluated in Equation (14), it is assumed that the same variation of Sdv with depth recurs at all other cross sections parallel to the section at the centre of the contact. However, in practice, contact pressure decreases at planes away from the centre of the contact, thus reducing the Sdv variation to lower magnitudes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000789_iros.2013.6696909-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000789_iros.2013.6696909-Figure5-1.png",
        "caption": "Fig. 5. Geometric model for describing the convex vehicle surface.",
        "texts": [
            " Using surface fitting techniques, the pressure readings from sensors at multiple locations can be used to reconstruct the pressure distribution over the entire body. Thus, the total damping force and moment exerted on the vehicle can be estimated by integrating the pressure distribution over the vehicle profile (excluding the locations of the actuators). The resultant force estimation will take the form of linear, fixed weight combinations of the pressure measurements. A. Vehicle Profile As illustrated in Fig. 5, the profile of the vehicle is defined in the body-fixed coordinate system. For a point on the vehicle\u2019s surface, the position vector r \u2208 E 3 and the normal vector n \u2208 E 3 can be written as r = (rx, ry, rz) , n = (nx, ny, nz) , (6) where rx, ry , and rz are components in r along the x-, y-, and z-axes, respectively; similarly for nx, ny, and nz . In general, the boundary of the vehicle can be represented with a biparametric surface, and thus, any position on the surface is uniquely defined by specifying a pair of parameters. In this paper, all extruding parts on the vehicle are ignored for simplicity. This gives a convex vehicle surface on which positions are conveniently determined by angles \u03b8 \u2208 [\u2212\u03c0, \u03c0) \u2286 R and \u03c6 \u2208 [\u2212\u03c0/2, \u03c0/2) \u2286 R, as illustrated in Fig. 5. As a result of the parameterization, vectors r and n are expressed as functions in \u03b8 and \u03c6. Suppose a number of p = p\u03b8 p\u03c6 \u2208 N sensors are located on the surface of the vehicle (p\u03b8 and p\u03c6 in each corresponding direction). Each sensor takes measurement of the normal pressure Ps \u2208 R at position \u03b8s, \u03c6s \u2208 R (s = 1, 2, . . . , p). Surface fitting over the pressure measurements will give an estimate of the pressure distribution P\u0302 (\u03b8, \u03c6) \u2208 R. A B-spline surface is used to model the pressure distribution due to its flexibility in the spline degree and smoothness, and its linear property that will become helpful for online computation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002612_dynamics.2016.7819064-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002612_dynamics.2016.7819064-Figure4-1.png",
        "caption": "Fig. 4. Angle Us projection obtained by vector Q values implementation from region Q for configurations: \u0430 qi (25\u043e, 20\u043e, 20\u043e, 25\u043e, 25\u043e); b qi (25\u043e, 20\u043e, 65\u043e, 25\u043e, 65\u043e).",
        "texts": [
            " notations \u041e12(1), \u041e12(2), \u041e12(3) accordingly define the position of the output link center (arm) in different planes of projections. Fig. 5a shows the results of android motions virtual modeling, where the criterion of motion volume minimization is applicable as well as forbidden regions are absent. Fig. 5b represents motions synthesis where a forbidden region is available, the developed algorithm of calculation generalized velocities vector Q values being used. Notations D\u043d (D\u043d(1), D\u043d(2), D\u043d(3)) \u0438 D\u0446 (D\u0446(1), D\u0446(2), D\u0446(3)) in Fig. 4 define the projections of the initial point and a goal under motions synthesis. Notations \u0420(1), \u0420(2) and \u0420(3) assign the position of the forbidden region on three projection planes. V. RESULTS DISCUSSION. SUMMARY AND CONCLUSIONS Relationships (2, 4, 5) when the region Q is assigned by the diamond permits the time of test assignment being reduced as opposed to the mode where k1 and k2 values are calculated based on a parallelepiped. The research results can be applied in intelligent control systems at the stage of robot arm virtual interaction with the external environment based on these models"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002091_cacs.2015.7465994-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002091_cacs.2015.7465994-Figure1-1.png",
        "caption": "Figure 1. Marine ship.",
        "texts": [],
        "surrounding_texts": [
            "Let us consider the mathematical model of the sea-going ship: u tdhbaa tdhbaa         , ),( ),( 222221 111211 (1) Here \u03c9 is an angular velocity relative to the vertical axis,  is a course (the turn to port side is considered positive),  is a deviation angle of the vertical rudders, \u03b2 is a drift angle (angle between the velocity vector and longitudinal axis of the ship), u is a control, d(t) is a bounded exogenous disturbance: u = k1 \u03b2+k2\u03c9+k3  +k4 \u03b4 (3) where k1, k2, k3, k4 are parameters need to be found those provide the desired dynamics of the closed-loop system. Deviation of the rudders and the its velocity turn (that is control) are constrained: \u2264 30o, u\u2264 3o/sec. Let us designate x = (\u03b2, \u03c9,  , \u03b4)T as a state vector, y(t) as an output vector. Then the system (1) may be rewritten as: , ),( Cxy tHdBuAxx   (4) where matrices A, B, C, D are:                                                     0 0 , 1000 0100 0010 0001 , 1 0 0 0 , 0000 0010 0 0 2 1 22221 11211 h h HCB baa baa A . Then the closed-loop system is modified to: , ),( Cxy tdHxAx cc   (5) where Ac = A + BKC = A + BK, Hc = H. For the construction of this controller let use the method, associated with evaluations of the reachability set, that is the range of output\u2019s values of the system which we can obtain. We have to determine the minimum ellipsoid, in a sense, comprising the optimized output y under any bounded disturbances d(t)."
        ]
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure10-1.png",
        "caption": "Fig. 10. Hypoid gear pair.",
        "texts": [
            "27 reduction ratio. A summary of the hypoid gear data is presented in Table 1 . It is common to define planar flanks in terms of a rack cutter or an involute profile where the resulting gear flanks are identical. Conjugate hypoid flanks were presented earlier in terms of rack cutters [2] and presented here in terms of involutoid profile segments. It is also common to reference axial and transverse displacements along with axial and transverse surfaces when discussing planar gearing. Presented in Fig. 10 is a virtual model of the automotive hypoid gear pair. A collage of images based on this gear pair is presented in Fig. 11 . These images include axial and transverse surfaces for hypoid gearing, the gear elements, and other combinations of gear elements and surfaces. The intersection between a gear element and a transverse surface is a transverse profile. A transverse profile is the focus of planar gearing. In general, the spatial involutoid curve is not a segment of a transverse profile. Presented in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003678_andescon50619.2020.9272048-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003678_andescon50619.2020.9272048-Figure1-1.png",
        "caption": "Fig. 1. Robot Bicycle Model",
        "texts": [
            " This paper is structure as follow: Section II presents the kinematic equations for the leader and follower robots which are four-wheel vehicles that will attain a formation. In Section III the linear approximation and I/O linearization control strategies are revised, also a collision avoidance algorithm is discussed. The proposed control methods are also presented. Section IV shows the simulation results. Finally, the paper ends in Section V with conclusions. This paper uses the Ackerman\u2019s Bicycle Model of a fourwheel mobile robot shown in Fig. 1. The bicycle model has a rear wheel fixed to the robot and the front wheel rotates In [1], Cao and his team give a critical survey of existing works and discuss open problems in the field. Based on that we know that there are different studies that can be found in the literature in relation with mobile robot formation: autonomous ground vehicles, unmanned aerial vehicles, autonomous underwater vehicles, and satellites. Various control strategies for mobile robots formation can be found in the literature",
            " This structure has a low manoeuvrability due to the Instantaneous Center of Rotation (ICR) and its nonholonomic restrain. However, it has an acceptable stability and a simple model which helps in the controllability. To find the dynamics of the bicycle model we need to assume some important restrains: \u2022 The mobile robot moves at low speeds. \u2022 The wheels do not side-slip while rotating. We also need to take into consideration the nonholonomic restrain which can be represented with (1) . . y cos(\u03c6)\u2212 . x sin(\u03c6) = 0 (1) Using geometric techniques on Fig. 1 and taking into consideration all of the restrains, we can get the state-space model given by (2) which describes the velocities and the turn rate of the vehicle. The parameters x and y are the coordinates of the rear wheel in the X-Y plane, \u03c6 is the angle of the mobile robot with respect to the X-axis, \u03b4 represents the steering angle, L is the length of the car and v the backward speed. Additionally, the leader robot will follow a path towards a desired final position with a final inclination represented by (x\u2217, y\u2217, \u03c6\u2217)T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002357_j.aej.2016.06.015-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002357_j.aej.2016.06.015-Figure1-1.png",
        "caption": "Figure 1 Schematic of compliant journal bearing.",
        "texts": [
            " Kuznetsov [19] has developed a numerical THD model to investigate the effect of lining compliance on the bearing characteristics. The analysis showed increased load carrying capacity, significantly reduced peak pressures and thicker oil film in the loaded zone compared to a white metal bearing. Slightly thinner oil films were predicted at the bearing edges. It was also shown that load carrying capacity was more sensitive to thermal expansion while pressure and oil film thickness profiles were more sensitive to elastic deformation. 2. Modeling of foil support structure The structure of bump foil bearing is shown in Fig. 1. It is comprised of a bearing sleeve lined with corrugated bumps (bump foil), the leading edges of both bump and top foil are ent of bump foil. f airfoil bearing based on bump foil structure, Alexandria Eng. J. (2016), http:// spot-welded to the bearing sleeve, the trailing edges of foils are free, and bump foils support the single flat top foil. The bump foils acting as springs provide stiffness, and the smooth top foil layer provides the bearing surface. When the shaft rotates over one certain speed, the top foil expands outward, and the air film is generated, and the shaft is then separated from the top foil"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001204_roman.2013.6628553-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001204_roman.2013.6628553-Figure8-1.png",
        "caption": "Fig. 8. Projection of the repulsive force generated on each spherical element on the manipulator arm to an equivalent force at the wrist/elbow.",
        "texts": [
            " When a sphere and a cylinder are in contact with each other, that is pij < (ri + rj), a virtual repulsive force for collision avoidance is generated between them. Referring to Figure 7, the direction of the virtual force is directed along the vector \u2212\u2212\u2192 PiPj and the magnitude of the force F is shown below: Fij = K|pij \u2212 (ri + rj)| (9) where K is a spring constant. During collision of robot segments, consider the case of a repulsive force Fr1 being generated on the upper arm at a distance of lr1 from the shoulder joint as shown in Figure 8. The force Fr1 can be transformed into an equivalent force at the elbow using the expression below: Felb = Lr1 L1 Fr1 (10) Similarly, a repulsive force Fr2 generated on the forearm at a distance of lr2 from the elbow joint can be transformed into an equivalent force at the wrist by: Fwr = Lr2 L2 Fr2 (11) For multiples collisions between bounding spheres, Felb and Fwr become a summation of the equivalent forces at the elbow and wrist respectively. As shown in Figure 9, the resultant force Felb produces a moment about the centre of the circle which traces all the possible elbow positions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002852_1045389x20906474-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002852_1045389x20906474-Figure8-1.png",
        "caption": "Figure 8. Flow characterization test setup. The red arrows represent the MRF path during flow tests. The rotor recirculation holes are blocked for this test.",
        "texts": [
            " The torque of the clutch is measured through its lever arm by a Transducer Techniques MLP-300 load cell, and it is controlled by a closed-loop controller. Clutch pressure is monitored by a 15-psi pressure transducer from SSI Technologies. Two thermocouples are located inside and outside of the shear zone in the magnetic circuit to monitor temperatures. During durability tests, the clutch temperature is maintained constant by water-cooling and an industrial chiller (Cornelius CH1500), insuring that the clutch maintains a temperature of 30 C regardless of power dissipation through the MRF. The flow characterization test setup is shown in Figure 8. An inlet tank containing Lord 140CG MRF is connected to the center of the clutch (instead of the expansion chamber) through a 9.5-mm inside diameter hose. The fluid flows from the inlet tank to the clutch. The rotor recirculation holes are blocked for this test, and the MRF is forced to travel radially from the center of the clutch to the shear zone by traveling at the right of the rotor. Then, it is pumped around the drum by the magnetic screw pump. Finally, it exits the clutch on the left side of the rotor through a second hose of 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000870_infcom.2013.6567020-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000870_infcom.2013.6567020-Figure1-1.png",
        "caption": "Fig. 1. The networks of (a) coal mine tunnels; (b) corridors of buildings; (c) underground tunnels. These three networks arc homotopically equivalent to (d) a 3D genus-2 network.",
        "texts": [
            ", INTRODUCTION Recent years have witnessed a rapid growth of Wireless Sensor Networks (WSNs) with generic tunnel-shape in emcrging applications where nodes are typically deployed on the surface of complex-connected 3D settings. Examples of such applications include monitoring of coal mine tunnels for disaster prevention and rescue,fire detection in the corridors of buildings.as well as monitoring of underground tunnels used in water, sewer or gas systems. These networks are called 3D high genus WSNs [22] or WSNs on high genus 3D surfaces, as shown in Fig. 1, The sensor network on a high genus 3D surface is often of a com pi ex-connected 3D setting and has non-trivial topology, possibly with high genus (i.e., multiple handles) [22], While there exist a series of previous studies This work was supported in part by the National Natural Science Foun dation of China under Grant 61073147. Grant 61173120. Grant 61103243. Grant 61202460. Grant 61271226. and Grant 61272410: by the National Natural Science Foundation of China and Microsoft Research Asia under Grant 60933012; by the Fundamental Research Funds lor the Central Univcrsirics under Grant 2012QN078: by the CHUT1AN Scholar Project of Hubci Province; by the Youth Chcnguang Project of Wuhan City under Grant 2O1O50231O8O; by the Scientific Research Foundation for the Relumed Overseas Chinese Scholars (State Education Ministry); by the National Natural Science Foundation of Hubei Province under Grant 201 ICDB044; by the Fok Ying Tung Education Foundation under Grant 132036; by the Hong Kong Scholars Program; and by the Program for New Century Excellent Talents in University under Grant NCET-10-408 (State Education Ministry)",
            " This section presents SINUS algorithm which deals with a fundamental problem: embedding a network on a high genus surface into a planar surface as a whole. The basic idea behind SINUS is simple: slicing the genus-H surface M to a simpler surface for embedding. The prcproccss of SINUS is to compute a triangulation from the original network via a simple distributed algorithm in [17], [23]. The triangulated structure, or mesh for short, forms a shape representation of the high genus surface, as shown in Fig. 1(d). Without leading to confusion, we still call the triangulated mesh as the high genus surface, denoted by M henceforth. Overall SINUS mainly consists of four steps: 1) We identify a maximum cut set Cmax from M, which is used to slice M to form a genus-0 surface G with 2n boundaries. To this end, a flooding across M is initiated by an arbitrary node r. This implicitly approximates the geodesic patterns on M and constructs a Morse function by assigning each node a level index. Based on the Morse function, a Reeb graph R starts at node T is also constructed in the process, by assigning each node a region ID"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001693_j.measurement.2015.05.016-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001693_j.measurement.2015.05.016-Figure6-1.png",
        "caption": "Fig. 6. Test platform of Cardan shaft dynamic unbalance.",
        "texts": [
            " SVD of E was carried out by applying Formula (6) and reconstruction of matrix E was carried out by applying Formula (8). Once that has been applied, the DDDR of vibration acceleration has been completed. The Fourier spectrum of DDDR is used to detect the fundamental, multiplier, and divider frequency caused by the cardan shaft dynamic unbalance. In order to validate the effectiveness of applying the DDDR model for the detection dynamic unbalance in Cardan shafts, the dynamic unbalance test bench (Length is 3.4 m, Width is 0.87 m) depicted in Fig. 6 was built. The power transmission path of the dynamic unbalance test bench is as follows: motor, gearbox, gimbal, cardan shaft and gimbal. The special revised cardan shaft which needs repair or replacement due to dynamic imbalance is installed on the test platform to be subjected to experimentation. The concrete testing work conditions of the experiments are listed in Table 1. The vertical acceleration on the gimbal next to the gearbox was collected at the sample rate of 20,000 Hz. A low-pass filter with a cutoff frequency of 2500 Hz was designed to filter experimental data as a preprocessing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure2-1.png",
        "caption": "Fig. 2 Standard imaginary generating gear",
        "texts": [
            " In this present study, both pinion and gear, being cut by the same planar-generating gear that results in conjugated tooth surfaces, are produced by the generating method (Fig. 1). Hence, if the cutters are arranged in the same position as the generating gear, they will produce a conjugated gear pair. Here, the generating gear is a planar gear with a pitch angle small than 90 deg, which gives the produced gear a little relief in the profile direction. In the standard planar-generating gear shown in Fig. 2, the right and left blades move along lines ocP and ocQ, respectively, which are located on the machine plane and past the apex point. The loci of the blades form the teeth of the generating gear. The parameters of the standard planar-generating gear are the pressure angle a0, the addendum at heel ha, and the thickness angle at tip aa. The position vector r \u00f0s\u00de c and surface normal vector n \u00f0s\u00de c of the tooth surface in the generating gear coordinate system S \u00f0s\u00de c can be represented by the following equation: r \u00f0s\u00de c \u00f0u;b\u00de \u00bc x \u00f0s\u00de c y \u00f0s\u00de c z \u00f0s\u00de c 1 h iT n \u00f0s\u00de c \u00f0u;b\u00de \u00bc @r \u00f0s\u00de c \u00f0u;b\u00de @u @r \u00f0s\u00de c \u00f0u;b\u00de @b @r \u00f0s\u00de c \u00f0u;b\u00de @u @r \u00f0s\u00de c \u00f0u;b\u00de @b 8>>>>>< >>>>: (1) where x \u00f0s\u00de c \u00f0u; b\u00de \u00bc b cos\u00f0aa=2\u00de y \u00f0s\u00de c \u00f0u; b\u00de \u00bc u cos a0 z \u00f0s\u00de c \u00f0u; b\u00de \u00bc 6\u00f0u sin a0 \u00fe b sin\u00f0aa=2\u00de\u00de 0 u 2:25mn; rt b rh 8>>>< >>: Here, u and b are the parameters along the tooth height direction and face width direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002511_j.jart.2016.09.006-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002511_j.jart.2016.09.006-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of the inverted cart-pendulum.",
        "texts": [
            " The origin of the boundary layer system (21) is exponentially stable \u2200Z. The origin of the reduced system (25) is exponentially stable. . Simulation results In this section, we will give two illustrative examples to show he applicability and efficiency of the proposed cascade control cheme. The first example is an inverted cart-pendulum system nd the second one is a ball and beam system. Controlling both ystems has practical importance. .1. Inverted cart-pendulum system Consider the familiar inverted cart-pendulum system (Aliddabi, 2005), depicted in Figure 2. The cart must be moved sing the force u so that the pendulum remains in the upright w osition as the cart tracks varying positions at the desired time. he differential equations describing the motion are (Al-hiddabi, 005): (Mp + m)y\u0308p + ml\u03b8\u0308 cos(\u03b8) + ml\u03b8\u03072 sin(\u03b8) = u l\u03b8\u0308 \u2212 y\u0308p cos(\u03b8) \u2212 g sin(\u03b8) = 0 (49) here \u03b8 is the angle of the pendulum, yp is the displacement of he cart, and u is the control force, parallel to the rail, applied o the cart. The numerical parameters of the inverted pendulum ystem are given in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000382_ijhvs.2018.089897-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000382_ijhvs.2018.089897-Figure7-1.png",
        "caption": "Figure 7 Variables of the mechanism position (see online version for colours)",
        "texts": [
            " 0$ or 0 , $i i iA A i i i s s s F M s \u00d7    = =       (2) where s0i is the instantaneous position vector of the reference point i related to the inertial reference point of the mechanism, si is the wrench orientation vector of the constraints i, Fi is the constraint force applied on joint i, and Mi is the constraint moment applied on joint i of the mechanism. In a more compact form, the wrench can be represented by equation (3). \u02c6$ $ .\u03a8A A= (3) where $\u0302A is the normalised wrench and \u03a8 is its magnitude. The proposed model (Figure 7) represents a trailer model making a turn. To simplify the model, the following considerations were made: \u2022 for the x direction, a steady-state model was used in the analysis \u2022 disturbances imposed by the road and lateral friction forces (Fy) (tyre-ground contact) in joints 3 and 19 were neglected \u2022 the total weight of the trailer (W), the inertial force (may), the tyre normal load FTi, the spring normal load FLSi, fifth-wheel normal load FFWi, and the passive torsional moment Txi are the only forces acting in the model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure10-1.png",
        "caption": "Fig. 10: The analysis procedure is improved making a preliminary TH simulation at the expected rotor frequency to get the actual bars current distribution to be implemented in the next MS simulations.",
        "texts": [
            " The consequence is also a better estimation of the rotor induced current for a certain slip and stator voltage supply. Carrying out the passages in Fig. 3, considering an uniform rotor current distribution, the field solution that satisfy the voltage condition is represented in Fig. 9. The procedure can be improved making a TH linear simulation to get the actual rotor bars current distribution. Both the stator and the rotor are supplied with equal and opposite currents in order to reproduce what happens along the q-axis. From the TH simulation field solution in Fig. 10a, the real part of the solution represents the time instant of interest to store the current distribution in each slots sectors. From the field solution in Fig. 10a the current in integrated in each slot regions and the ratio with the total bar current gives the part of the bar current that flows in a certain slot sector. The current in p.u. in the i-th sector of the j-th bar is Authorized licensed use limited to: Carleton University. Downloaded on June 30,2020 at 13:07:46 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (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. (a) TH simulation performed to get the rotor bars current distribution. Stator and rotor current are imposed equal and opposite to replicate what happens along the q-axis; then the real part of the solution represents the time instant of interest. obtained as: iij = < ( ip,i ) < ( ibar,j ) (9) The final matrix represents the subdivision of the bars current in the several sectors that has to be implemented in the static analysis. The final result is shown in Fig. 10b. The solution is achieved iterating the procedure in Fig. 3, until the current isd and isq satisfy the voltage constraint. The obtained stator and rotor current will carefully take into account not only the iron saturation, but also the actual value of the rotor resistance and leakage inductance, that directly depend upon the bars current distribution. The LR operation is a particularly critical working point, since the stator and rotor current (and the IM starting torque) closely depends on an accurate knowledge of the rotor parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001811_s11431-015-5883-3-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001811_s11431-015-5883-3-Figure5-1.png",
        "caption": "Figure 5 (Color online) Sketch of SFB [44].",
        "texts": [
            " [28] observed the film thickness deviates from the Dowson- Higginson equation varied with speed, as well as the time effect on the film thickness, and the failure condition of the thin film lubrication [40]. The surface force apparatus (SFA)/Surface force balance (SFB) was initially developed for measuring the interacting force, including both normal and lateral forces, between two solid surfaces across nano/sub-nano gap. Based on the fringes of equal chromatic order (FECO) technique, the device can resolve distances to within 0.1 nanometer. The substrate for the surfaces is generally atomically smooth mica, which can be moved towards and retracted from one another. Figure 5 shows the surface force balance schematically. Two pieces of single facet mica with atom-scale roughness, which were back-silvered, are usually glued down onto cylindrical lenses and then subjected to the experiment. The interaction geometry can be also determined from the shape of interference fringes. Normal forces, Fn, are directly inferred as a function of the separation by monitoring the increments of the normal force obtained from the bending of the spring. The shear forces, Fs, are determined from the bending of lateral spring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000274_acemp-optim44294.2019.9007164-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000274_acemp-optim44294.2019.9007164-Figure3-1.png",
        "caption": "Fig. 3. (a) Image of monolithic TX with stacked CMOS drive and periodic loaded MZM. The MZM has a total push-pull RF/optic interaction length of 6 mm (3 mm per MZM arm). (b) Magnified view of periodically loaded electrode in MZM.",
        "texts": [
            " An MZM electro-optic modulator is used in the TX due to its relatively temperature insensitive operation. The MZM has lateral PN junctions with 4 \u00d7 1017 cm\u22123 peak p and n doping concentrations (as indicated by simulations of the processing implant conditions and thermal budget), a \u223c135 nm SOI thickness, and 2 \u03bcm BOX appropriate for a \u223c1.3 \u03bcm wavelength of operation. The pn junctions show a leakage current on the order of nanoamps with a 1 V reverse bias, and so have a shunt resistance on the order of hundreds of mega-ohms. An image of the TX is shown in Fig. 3(a). The CMOS driver differential output is coupled directly into a periodically loaded push-pull MZM. The MZM incorporates unloaded RF electrode segments to increase the electrode effective line impedance. The unloaded RF segments are transmission line sections that are not connected to pn junctions and are periodically inserted into the MZM electrode. Fig. 3(b) shows the loaded sections (electrode sections coupled to a pn-junction) are the straight horizontal electrode segments highlighted by thick white arrows, and the unloaded sections are the 180\u00b0 turn segments denoted by the narrow white arrows. Each MZM arm has ten 300 \u03bcm long electrode sections loaded with optical pn-junctions (3 mm total loaded electrode length in each MZM arm) and nine 157 \u03bcm long unloaded elec- trode sections, resulting in a \u223c68% loading and 4.41 mm total electrode length. Since the unloaded electrode sections have low capacitance their relative contributions to RF propagation losses are small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.58-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.58-1.png",
        "caption": "Figure 3.58 Spur gearhead schematic",
        "texts": [
            " \u2022 Synonymous with this, the load inertia seen at the gearhead input is the load inertia divided by the square of the gear ratio, resulting in a close match to the motor inertia (ideally 1:1), creating the highest overall system efficiency (see Section 4.5). Although available in many configurations and technologies, gearheads using the following are those most frequently found in high performance, closed loop servo systems: \u2022 Spur gearing \u2022 Planetary gearing \u2022 Hybrid gearing \u2022 Worm gearing \u2022 Harmonic gearing. The spur gear is the oldest design, going back to as early as 30 BC and from which all other configurations have evolved. A simple schematic for a one pass spur gear arrangement is shown in Figure 3.58. The smaller gear, called a pinion, is usually mounted on the input shaft and the larger gear is connected to the output shaft. A single pass, as shown, is usually applicable to no more then a 10 or 15 to 1 ratio. Larger ratios are accomplished by using two or three passes, with the pinion of succeeding passes mounted on the same shaft as the larger gear of preceding passes. 108 Electromechanical Motion Systems: Design and Simulation System Components 109 One deficiency of the basic spur gear design is that in its simplest configuration, as shown, the input shaft and output shaft are not in line"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure6-1.png",
        "caption": "Fig. 6. Proposed 6S-13P E-Core CPSFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " Analysis reveals that conventional 6S-10P E-Core SFPMM and 6S-10P C-Core SFPMM retain same PM volume as that of 12S-10P E-Core SFPMM and increased slot area, however author fails to compensate effects of leakage flux. Until now, many researchers tried to reduced PM as much as possible and suppress flux leakage but unfortunately both effects are not considered at a time. In this paper, alternate Consequent Pole SFPMM (CPSFPMM) with partitioned PMs are introduced which further reduce PM usage and suppress flux leakage completely. For analysis and electromagnetic performance evaluation, this paper proposed three topologies of CPSFPMM (as shown in Fig. 5, Fig. 6, Fig. 7) corresponding to three topologies of conventional SFPMM as listed in table I. Based on Finite Element Analysis (FEA) proposed model has successfully reduced PM usage much more and suppress PM leakages from the end and enhance flux modulation effect. Hence proposed model reduces machine cost furthermore, retaining electromagnetic performance. The rest of the paper is organized as, section II present design parameter, design methodology and working principle of proposed CPSFPMM, section III illustrates FEA based electromagnetic performance analysis and finally, some conclusions are drawn in section IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003332_aim43001.2020.9158809-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003332_aim43001.2020.9158809-Figure6-1.png",
        "caption": "Figure 6. The velocity potential field Hstatic and Hdynamic",
        "texts": [
            " (15) is calculated by: , , , k i i L k i k k k i k i k i i O R R O R O MO R O M R O M MR R  \u2208\u0399  \u22c5= \u2212 \u2208\u0399\u0399   \u2208\u0399\u0399\u0399 o        (16) The stretch vector kO M  is defined as follows: 2 O k OO M t \u03b2 = \u22c5\u2206 \u22c5 v v  (17) where 0\u03b2 > is a coefficient to adjust the range of the dynamic obstacle and t\u2206 is the interval time. As the obstacle speed Ov increases, the stretch vector kO M  increases at the same time. This means that the range of dynamic obstacles will increase. The velocity potential field staticH and dynamicH is shown in figure 6. III. FORMATION CONTROL WITH DYNAMIC OBSTACLES As mentioned in Section II, the control velocity of the robot is composed of three parts: 1)the velocity conv used to form and maintain the formation under the consensus protocol; 2)the repulsion velocity RRv between robots by constructing the velocity potential field; 3)rejection velocity ROv generated by obstacles. The velocity composition realizes the flexible switching of the formation maintenance and obstacle avoidance for each vehicle. The total speed is obtained by the following equations: consum RR RO= + +v v v v (18) max max max , if , if sum sum sum sum sum v v v  \u2264 =  >  v v v v v v (19) IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure3-1.png",
        "caption": "Fig. 3. The coordinate systems of a multiple-mode equivalent 8R mechanism.",
        "texts": [
            " When the antiparallelogram units deform, the whole mechanism will change its config- uration correspondingly. To simplify the kinematic analysis, we can regard the multiple-mode mechanism as a spatial 8R mechanism with variable link lengths as shown in Fig. 2 (b), the link lengths changes according to the movement of the antiparallelogram unit, that is, different configurations of the multiple-mode mechanism correspond to different spatial 8R mechanisms. The coordinate systems are set up in the multiple-mode mechanism, as shown in Fig. 3 . Let a j be the length of the common normal distance from z 2 j -1 to z 2 j , R 2 j + 1 be the common normal distance from x 2 j + 1 to x 2 j + 2 , and \u03b82 j be the rotation angle from x 2 j to x 2 j + 1 . According to the parameters and motion characteristics of the antiparallelogram unit, these D-H parameters can be calculated as: where j = 1,2,3,4, R 9 = R 1 and \u03b8A 5 = \u03b8A 1 . l denotes the lengths of A j D j and B j C j of Fig. 2 (a), l 1 denotes the lengths of A j B j and C j D j of Fig. 2 (a), and l 1 > l + d ( d is the diameter of revolute joint)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure1-1.png",
        "caption": "Fig. 1. Design of a 3-DoF TPM (1. Fixed base, 2. Rotary joint, 3. Planar joint, 4. Active prismatic joints).",
        "texts": [
            " The hydraulically actuated 3-DoF TPM was developed and tested for the steel industry's needs. For this reason, the main goal of this work was to test IEHSDs in controlling the trajectory tracking of the 3-DoF TPM for handling heavy forgings. A prototype of hydraulically actuated 3-Dof TPM with IEHSDs was constructed and tested in the Department of Mechatronic Devices (Kielce University of Technology, Poland). The TPM prototype actuated by three IEHSDs with construction details is shown in the Fig. 1. The TPM has a fixed base, movable platform and three IEHSDs (servohydraulic axes). The hydraulic cylinders are mounted with rotary joints to a fixed base on the circumference of R = 0.25 m and the piston rods are mounted with planar joints to the moving platform on the circumference of r = 0.130 m. Each single integrated IEHSD (Bosch Rexroth) consists of a hydraulic actuator (piston diameter D = 0.040 m, rod diameter d = 0.028 m, stroke h = 0.250 m, nominal pressure p = 16 MPa), proportional directional control valve (nominal flow qvn = 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure7-1.png",
        "caption": "Fig. 7 Different kinematic schemes with the same structural scheme",
        "texts": [
            " It can be determine as follows: The direction of the circulating power Pcirc coincides with the direction of input power PA or output power PB at the coupling shaft (constituting the external compound shaft) where its algebraic sign of the internal torque coincides with the algebraic sign of the torque of external compound shaft. The directions of the relative powers Prel I and Prel II in both component trains that are important for the determination of the efficiency are shown also. In the gear trains with internal power circuits it is shown that with one and the same structural scheme, the kinematic schemes may differ (Fig. 7). Figures 10 and 11 show two examples of the compound two-carrier planetary gear trains with two compound shafts and four external shafts. The brakes used in these gear trains can be located on different shafts\u2014single or compound. The Fig. 10 shows a reverse gear train, carrying forward and backward. The Fig. 11 shows a change-gear (gearbox), carrying out two gear-ratio steps (speeds). Figure 12 shows three cases of using two-carrier compound planetary gear trains with a one compound shaft, but with four external single (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002626_iccas.2013.6704133-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002626_iccas.2013.6704133-Figure4-1.png",
        "caption": "Fig. 4 Measurement between a mobile robot and a point landmark.",
        "texts": [
            " = O. It is obvious that if two lines are parallel, the state is unobservable and cannot be estimated. In Fig. 3, the condition number is an infinite value when two lines are parallel, so it is unobservable. In addition, if the lines are near parallel, the condition number increases drastically. By contrast, condition number is the smallest when the two lines are orthogonal and the robot can estimate its state in this situation. In the relationship between a robot and a point land mark shown in Fig. 4, the measurement model is ex pressed as (9): z = h(x) = [R] = [J \ufffdX2,: \ufffdy2 ] B arctan \ufffd - 1jJ (9) where \ufffdx = XR - XL and \ufffdy = YR - YL. The observability matrix of point landmarks is more complex than the case involving lines. Furthermore, the analytic solution relative to the condition number is hard to be derived exactly for the point landmarks. To obtain an intuitive understanding and to access the phenomena easily, we assume the situation with v = 0 and U! = 0 and instead of selecting the method for solving analytic equa tions we use a method that involves substituting the land mark information into the observability matrix directly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000741_jfm.2014.45-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000741_jfm.2014.45-Figure5-1.png",
        "caption": "FIGURE 5. (Colour online) The organized flight of geese. Around the right (or left) trailing goose, the black curves with arrows are the streamlines of the virtual velocity \u2207\u03c6T (or \u2207\u03c6L) caused by the virtual goose moving along the thrust (or lift) direction.",
        "texts": [
            " In this section, the mechanism of the V-shaped formulation is discussed in order to find the optimal angle where the trailing goose experiences high thrust or lift. According to the vortex contribution T4 in (3.25), the distributed vorticity in the flow can affect the aerodynamic forces on the adjacent geese (ahead/behind) in the formation flight, i.e. T4 =\u2212 \u222b b 0 \u03c1 [\u0393 \u00d7 u] \u00b7 \u2207\u03d5 dx, (4.3) where the Lamb vector (\u0393 \u00d7u) is an objective existence, and the virtual velocity (\u2207\u03d5) needs to be determined. Suppose that the position (distance and azimuthal angle) is sustained in the V-formation flight of the geese (as shown in figure 5), and the virtual velocity caused by a complex goose body can be modelled as a sphere in the analysis, which is reasonable when the distance between two geese is large enough. First, we consider the thrust on the right rear goose moving virtually along the thrust (forward) direction. Here T4 is employed to predict the contribution of the leading goose\u2019s wake to the thrust. The virtual velocity potential excited by the forward virtual moving \u2018goose\u2019 (right-rear in figure 5) approximates to \u03d5T =\u2212u\u22170 a3 2r2 cos \u03b8T, (4.4) where a is the radius of the sphere, r is the radial distance from the centre of the goose\u2019s body to the point we considered and \u03b8T is the azimuthal angle between the zenith (forward) direction and the vector r (see figure 6). Then, the magnitude of the spanwise component of the virtual velocity is v\u2217 =\u22133u\u22170 2 (a r )3 sin \u03b8T cos \u03b8T, (4.5) which is located in the horizontal plane. In the above formula, \u2018\u2212\u2019 and \u2018+\u2019 indicate the left and right front regions of the goose, respectively",
            " It reveals that the leading goose experiences larger drag than it flies solely. According to (4.7), the drag reduction decreases with the distance d. The distance from the leading goose to the wake of the trailing one is larger than that from its wake to the trailing. So the increased drag on the leading goose is less than the drag reduction on the trailing one. Compared with the simple sum of the drag in solo flight, the total drag of the flock is reduced. Second, we consider the lift on the left-rear goose (left-bottom in figure 5), let it move virtually along the vertical direction, and generates the virtual potential \u03d5L. The streamlines are shown in figure 5. If the formation flight is in the horizontal plane, the magnitude of the spanwise component of the virtual velocity is zero, v\u2217 = 0, which results in T4 = 0. This means that there is no added lift for the trailing goose flying near the wake of the leading goose. In short, the V formation flight may benefit the geese when they migrate a long distance from one place to another in the view of aerodynamics. There is an advantage of drag reduction for the trailing goose while the increased drag happens to the lead one, but the total drag is decreased for the flock"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002601_humanoids.2016.7803410-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002601_humanoids.2016.7803410-Figure4-1.png",
        "caption": "Fig. 4: Foot posture deviation cost with CoM position. The left foot is stationary while the right foot is movable. The support polygon is shown in red with the red dot marking the center of mass projection. The yellow dots mark the center of the foot whose frames are written as FL and FR. FR\u2032 represents the right foot target frame.",
        "texts": [
            " The collision avoidance constraint limits the distances between all of the collision pairs to be greater than a tolerance threshold. By modeling the robot and the environment as a group of convex shape objects, the distance between two objects can be efficiently computed using the GilbertJohnsonKeerthi algorithm (GJK) [14]. We use a hinge-loss function introduced in [12] to set up the constraint. \u2022 Center of mass (CoM) constraint. The horizontal projection of the CoM has to be on the support polygon between two feet (see in Fig. 4). Similar to the second type Cartesian posture constraint, the support polygon can be represented as an intersection of k half planes with the form, a j(q)xCoM(q)+b j(q)yCoM(q)+ c j(q)\u2264 0, j = 1,2,3, ...,k where a j, b j, and c j are constants when both feet are fixed at the same position. When the feet are repositionable, the shape of the support polygon would change, and a j, b j, and c j can be computed given the variable q. xCoM and yCoM are the location of the CoM projection on the ground. Therefore, to generate a feasible IK solution for Valkyrie, a variety of costs and constraints need to be set, such as, joint displacement costs using a normal standing pose as the nominal pose, pelvis height and orientation costs, torso orientation costs, joint limit constraints, collision avoidance constraints, and CoM constraints",
            " 1) Fixing both feet: Add the first type Cartesian posture constraints on both feet. The desired postures of both feet are their current poses. 2) Fixing one foot: Add a second type Cartesian posture constraint for the movable foot, whose details are described in Section II-A. Add a first type Cartesian posture constraint on the stationary foot. Since the ideal position of the CoM projection is the midpoint of the line segment between the centers of the two feet, given the stationary foot frame and the CoM projection which are FL and (xCoM,yCoM) in Fig. 4, the desired movable foot frame FR\u2032 can be computed. Then a Cartesian deviation cost with a desired posture obtained from FR\u2032 can be added into the objective. The foot which has a shorter distance to the end effector target pose would be chosen as the movable foot. To prevent from generating an unnecessarily large step, another Cartesian deviation cost is added to penalize the movement of the movable foot from its current posture. 3) Relaxing both feet: Due to both feet being allowed to move in the x-y plane and rotate in yaw direction freely, but keep z position, roll and pitch orientation the same as their current configuration, the first type Cartesian posture constraints would be added on both feet with a diagonal matrix whose coefficients are [0,0,1,1,1,0]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001663_1.4024235-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001663_1.4024235-Figure1-1.png",
        "caption": "Fig. 1 Structural synthesis: (a) the constraining limb and (b) the CAD model of the 2 UPS 1 PRP",
        "texts": [
            " Section 2 treats of the structural synthesis of the mechanism, while other important issues such as the position analysis, kinematic errors, geometric errors and parameter sensitivity analysis are presented in Secs. 3\u20136, respectively. The structural synthesis goal is to generate a parallel mechanism, able to position the cutting tool in the 3D-space, in such a way that it performs only three independent motions. In this work, an alternative type synthesis procedure, proposed by Hess-Coelho [14,25], is employed. To constrain the tool motions, we choose an active kinematic chain PRP, connected to the mechanism platform (Fig. 1(a)). In order to describe the motion of the cutting tool, two reference frames are chosen: one in the fixed base O xb yb zb and another in the moving platform P xp yp zp. The axes of the revolute and the first prismatic joints are parallel to the vertical axis zb, while the axis of the second prismatic joint is orthogonal to the previous ones. Let the kinematic bond [26] between the base b and the moving platform p, associated to the passive limb I, is labeled as LI b;p . In accordance with the Lie\u2019s theory of continuous groups [12], the limb I is a generator of the cylindrical and prismatic displacement groups, along axes zb and xp, respectively LI b;p \u00bc C zb\u00f0 \u00de [ P xp (1) In order to control the motion of the moving platform, by employing the topology of a parallel architecture, we add to the original kinematic structure two active peripheral limbs, named as II and III. They correspond to generators of the general sixdimensional group of rigid-body displacements D LII b;p \u00bc LIII b;p \u00bc D (2) A feasible choice for each peripheral limb is the utilization of a UPS chain. Then, one possible mechanism for the task is the 2UPS\u00fePRP (Fig. 1(b)). One can notice that the set of displacements common to the three subgroups mentioned above, LI\u00f0b;p\u00de \\ LII\u00f0b;p\u00de \\ LIII\u00f0b;p\u00de, coincides with the group generated by the limb I. Therefore, the dimension of the resultant group of displacement is three, which equals to the connectivity between the base and the moving platform, enabling the cutting tool to perform three independent motions LI b;p \\ LII b;p \\ LIII b;p \u00bc C zb\u00f0 \u00de [ P xp (3) Regarding the actuation scheme, we set active one prismatic joint in each limb, resulting the 2UPS \u00fe PRP parallel mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003115_physrevfluids.5.053103-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003115_physrevfluids.5.053103-Figure6-1.png",
        "caption": "FIG. 6. Tracers\u2019 trajectories compared to rigid particles\u2019 trajectories after one beating cycle. Tracers\u2019 initial and ending positions are shown in black and red crosses, respectively; particles\u2019 initial and ending positions are shown in open and solid circles. The trajectories of the tracers and the particle centers are shown in solid and dashed lines, respectively. (a)\u2013(c) Particle radius is rp = 0.1; (d)\u2013(f) particle radius is rp = 0.4. Left to right: Nw = 0, 1, \u22129.",
        "texts": [
            " 5(b), at each of the humps on the outer wall, a shear region could be observed which does not exist in the case of the regular Taylor-Couette geometry. In this section we study the full cilia-channel-particle problem and compare the results with passive tracers, in an effort to showcase the effects of the particle size in such problems. We start by uniformly seeding 20 circular particles of radius rp inside the channel and trace their centroids within one ciliary beating cycle. With small particle size, as shown in the top row of Fig. 6, the differences between the passive tracers and the finite-size particles are hardly visible, as 053103-11 expected. With large particle size, however, the difference becomes much more evident. Specifically, in the first two phase differences ( \u03c6 = 0, \u03c0/16), the motions of the large particles are close to those of the passive tracers, albeit having a noticeably shorter distance traveled (dashed curves have 053103-12 shorter lengths compared to solid curves). In the case of large phase difference ( \u03c6 = \u22129\u03c0/16), the difference between the trajectories of the large particles and the tracers is even more evident"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000513_0954407015590703-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000513_0954407015590703-Figure15-1.png",
        "caption": "Figure 15. Tester machine for measurement of the relaxation length.",
        "texts": [
            " These small differences are related to the effect of the self-aligning torque, because they do not appear if the self-aligning torque is introduced into the first-order model. The effects of the tyre damping properties appear to be very small; this parameter influences strongly the modulus and the phase of the FRF only if it assumes very large values. Comparison with measured values In the framework of this research, experimental tests were performed on the rotating-disc tester machine at Padova University.10,17,18 This particular machine is equipped with a disc (diameter, 3m) that rotates about a vertical axis (Figure 15). The wheel under test rolls on the outermost part of the disc, which is covered by high-friction material to mimic the friction between the tarmac and the wheel in normal road conditions. The wheel is held in position on the rolling track by means of an hinged arm, which makes it possible to set the side-slip angle and the camber angle at assigned values. A brushless servomotor drives the camber rotation of the whole arm; the range is 654 . Another brushless servomotor drives the yaw rotation of the fork and its range is 610 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001284_j.ijrmms.2013.04.002-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001284_j.ijrmms.2013.04.002-Figure5-1.png",
        "caption": "Fig. 5. Basic concept of scaling and definition of unknown fragment size using principles of optics.",
        "texts": [
            " Because of that it was necessary to find a scaling method. Once the basic understanding of imaging was acquired the solution became quite simple. If the image contained the object of known dimensions (machinery for example or spacing/burden) and if the distance from the camera to the object was known it was possible to scale the dimensions of the object of interest knowing its distance from the camera and its size on the image by applying proportions (direct and inverse). The procedure is shown in Fig. 5. An object of known dimension AB is located at distance l1 from the camera lens. At the same time, its dimension in the image is A\u2032 B\u2032. Knowing these dimensions it becomes possible to calculate (pseudo) focal length f from the ratio AB : A\u2032B\u2032\u00bc l1 : f \u00f011\u00de f \u00bc A\u2032B\u2032\u22c5l1 AB \u00f012\u00de Now, if there is an object of unknown dimensions CD, located at known distance l2, and with known dimensions C\u2032D\u2032 in the image (measured), the dimension CD can be calculated from the ratio CD : C\u2032D\u2032\u00bc l2 : f \u00f013\u00de CD\u00bc C 0D0\u22c5l2 f \u00f014\u00de For example, when footage from April 21st 2011 was analyzed a suitable falling fragment was noticed at 1 min 07 s (frame \u266f2005)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.61-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.61-1.png",
        "caption": "Figure 3.61 Harmonic gearhead cross-section",
        "texts": [
            " This arrangement is unidirectional in that the worm cannot be driven by the worm gear and actually will bind and result in a braking action if the load attempts to back drive. Worm gearing has traditionally been used for high power, unirotational applications to obtain right angle, high reduction ratios in a compact assembly in which efficiency is of secondary concern. Contemporary improvements in tooth design, materials, lubrication, bearings and thermal design have created worm gear assemblies that are compatible with the severe requirements of closed loop servo operation. The harmonic gear is a unique assembly of three main components, as shown in Figure 3.61. The fixed ring gear has internal teeth. The flexspline is a flexible cylindrical member whose outer diameter is slightly smaller than the inner diameter of the fixed ring gear, has two fewer teeth than the fixed ring gear and is held in contact with the ring gear in only two areas 180\u25e6 apart, as determined by the elliptically shaped wave generator. As the wave generator rotates, the contact areas between the flexspline and the fixed ring gear rotate, such that when the wave generator rotates 360\u25e6 the contact area only moves two System Components 111 teeth from the starting position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000716_ecce.2015.7309909-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000716_ecce.2015.7309909-Figure4-1.png",
        "caption": "Fig. 4: Transformation from stator to rotor reference frame.",
        "texts": [
            " The magnetic potential of the rotor Ur(\u03b8r) is computed by means of the magnetic circuit of the machine, where the flux barriers are modeled as constant reluctances [1]. Fig. 3 shows the cross section of an eccentric rotor. An SPM machine is kept as an example. The point of symmetry of the rotor Or is shifted from the point of symmetry of the stator Os, by distance e and angle \u03b8e. To the aim of considering the air gap length variation, as shown in Fig. 3, the constant value of the inner stator radius (i.e., Rs) is replaced by a variable radius computed according to the rotor symmetric point Or (i.e., Rsr) [10]. As shown in Fig. 4, the cartesian coordinates of the point P in the stator reference frame are Pxs = Rs cos(\u03b8s) Pys = Rs sin(\u03b8s) (5) The transformation of the cartesian coordinates of point P from the stator to the rotor reference frame are as follows:[ P \u2032 xr P \u2032 yr ] = [ cos(\u03b8e) sin(\u03b8e) \u2212 sin(\u03b8e) cos(\u03b8e) ] \u00b7 [ Pxs Pys ] (6) and then Pxr = P \u2032 xr \u2212 e Pyr = P \u2032 yr (7) Then, Rsr and \u03b8sr are given by: Rsr = \u221a P 2 xr + P 2 yr \u03b8sr = tan\u22121(Pyr/Pxr) (8) The non-uniform air gap length can be approximated by: g(\u03b8r) = g0 \u2212 e \u00b7 cos(\u03b8r \u2212 \u03b8e) (9) 1) Eccentricity with SPM machine: Referring to the SPM machine, as shown in [10], Rsr and \u03b8sr change the coefficients anI , bnI , cnI and dnI in [2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001728_j.jappmathmech.2015.09.002-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001728_j.jappmathmech.2015.09.002-Figure6-1.png",
        "caption": "Fig. 6.",
        "texts": [
            "14) If the system begins to move as a consequence of the application of a certain additional force F, then according to the work-energy heorem ence it follows that (F, dq) > 0, which proves the Hill stability. This assertion has been proven. emma 8. The satisfaction of inequality (4.14) for any real displacements is sufficient for Hill stability. If all the normal reactions at the quilibrium position of a scleronomic system are non-zero and any virtual displacement results in the appearance of sliding friction (the eal displacements then appear in the set of virtual displacements), Hill stability occurs. xample 8. Consider a sphere clamped between two rough surfaces (Fig. 6). The equilibrium conditions are as follows: his system has a solution if (4.15) ince the sphere can simply fall down (in this case the reactions are equal to zero), inequality (4.15) ensures possible (not necessary) quilibrium. According to Lemma 5, such an equilibrium position is Lyapunov unstable, and according to Lemma 8, it is Hill stable. emma 9. Suppose all the reactions of the frictional constraints in an equilibrium position lie in the cone of friction, i.e., inequality (2.3) s strict. If for all the kinematically possible displacements dq that are not associated with the appearance of sliding friction the work of he specified generalized forces is non-positive, that is, if ill stability occurs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002312_1464419316661969-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002312_1464419316661969-Figure2-1.png",
        "caption": "Figure 2. Interaction between the ball and cage.",
        "texts": [
            "comDownloaded from moment vector Ma b is the same as gyroscopic moment, which makes the ball roll along the axis of the bearing and changes contact angle. The same method is used to calculate Fc rj and Mc rj, and the net force vector and moment vector acting on the raceway in the ball azimuth frame is given as Fi r \u00bc PNb j\u00bc1 TciF c rj Mr r \u00bc PNb j\u00bc1 TcrM c rj 8>>>< >>>: \u00f012\u00de where Tci and Tcr are the transformations from contact frame to inertial frame and race-fixed frame. Interactions between the ball and cage. As shown in Figure 2, geometrical interactions between the ball and cage are described. In this figure, cage-fixed frame CXcYcZc is used to define the cage centre in inertial frame, pocket frame DXdYdZd is used to confirm the pocket centre in the cage-fixed frame. The vector ric is expressed as the position of the cage centre C in the inertial frame, and the vector rcdc is used to describe the location of pocket centre D in the cage-fixed frame. Then, the vector between the ball centre and pocket centre in the cage-fixed frame can be written as rdbd \u00bc Tcd\u00f0Tic rib ric rcdc\u00de \u00f013\u00de where Tcd and Tic are the transformations from cagefixed frame and inertial frame to pocket frame and cage-fixed frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002827_j.optlaseng.2020.106065-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002827_j.optlaseng.2020.106065-Figure13-1.png",
        "caption": "Fig. 13. The state of the particle and the surrounding fluids at 12 \ud835\udf07s: (a) where the powder particle is located; (b) speed distribution of the particle and the surrounding fluids; (c) the speed distribution of the pool around the particle.",
        "texts": [
            " As an be seen from the figure, when the particle starts to enter the melt ool, the speed of the particle did not change substantially, which is 2.25 m s \u2212 1 . The gas-liquid interface remains basically horizontal, and he surface of the melt pool is not significantly disturbed, as the flucuation area on the gas-liquid interface is less than twice the cross secion of the particle. However, the melt pool is subject to a relative high peed and speed gradient around the particle, as the peak speed reaches 2.25 m s \u2212 1 . The powder particle continues to move into the melt pool via inertia t a movement time of 12 \ud835\udf07s to the position shown in Fig. 13 (a). Durng this time, the powder particle goes into the melt pool to a depth f about one third of its own diameter at a speed of ~2.03 m s \u2212 1 , hich is ~81.2% of its initial speed (combing with Fig. 13 (b), which hows the speed distribution of the particle and the surrounding fluids). ig. 13 (c) is the speed distribution of the melt pool around the partile, which demonstrates that the peak speed of the melt pool decreases o ~2.03 m s \u2212 1 , and the fluctuation area at the gas-liquid interface of he melt pool around the particle increases to more than twice the cross ection of the particle. When the powder particle moves to the position shown in Fig. 14 at moving time of 18 \ud835\udf07s and has an immersion depth about one-half of ts own diameter in the melt pool, the powder particle speed decreases o ~1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure1-1.png",
        "caption": "Fig. 1. Three-link direct-contact spatial mechanism with higher pair in point contact.",
        "texts": [
            " Well established analytical methods are developed for kinematic analysis of mechanisms containing only lower kinematic pairs [1]. Prominent among them are the Chase's vector loop [2] and the Denavit\u2013Hartenberg matrix methods [3]. But for mechanisms involving higher kinematic pairs, there do not exist any such general structuredmethods. In literature, kinematic analysis of mechanisms having higher kinematic pairs such as sphere and plane in contact [4], sphere in a cylindrical groove [4,5] is done using vectormethods. In certain direct-contact spatialmechanisms (for example see Fig. 1 and also [6]) and multi-body systemsoneoften encounters thehigher kinematic pair consisting of two general surfaces in point contact. For suchmechanisms above methods cannot be applied for kinematic analysis. A three-link spatial mechanism which includes one higher kinematic pair of two tori in point contact is analyzed using 7R (revolute) equivalent mechanism which is valid over the entire range of motion of the higher pair mechanism [7]. Generally in the analysis of a higher pair mechanism, one is often interested in the velocities and accelerations of the constituent links rather than the actual details of the relativemotion between the elements of the constituent higher pair joints",
            "\u03c9 l and \u03bc l be the angular velocity and acceleration of the constraining link-l respectively. Let r be the position vector of Kw.r.t. Ko. The U-joint at Ko prevents the spin motion of link-l in the direction of r . Then, \u03c9 l and \u03bc l can be uniquely solved for from the following equations: \u03c9 l r \u00bc vK=Ko \u00bc vK\u2212vKo \u00f02\u00de \u03c9 l \u03c9 l r\u00f0 \u00de \u00fe \u03bc l r \u00bc aK=Ko \u00bc aK\u2212aKo : \u00f03\u00de In addition, equations for velocity and acceleration of points K and Ko are obtained through constraining of the elements of the higher pair (for example, see the revolute joint constraints in Fig. 1). These set of equations can be solved for complete velocity and acceleration analysis of the mechanism containing this higher pair. The planar analogue of the virtual sphere is the osculating circle. In Fig. 5(a),m-curve and f-curve are the planar profile curves of the contacting rigid bodies in point contact. Cm and Cf are the centers of curvature ofm-curve and f-curve respectively at the point of contact C. For any relativemotionup to second-order, the center of curvature of the instantaneous point-path ofCm is located at Cf",
            " Using these coordinates, the two new substitution centers are chosen by means of Eq. (60). A value of z is chosen arbitrarily, then the corresponding zowill be determined from that equation. After the substitution, the acceleration analysis of themechanism is done using the existing methods for lower pair mechanisms. But it should be reminded that before this second step of acceleration analysis, velocity analysis of only the link-lmay need to be reworked since the joint centers of this link changed. Fig. 9 shows an illustration of the mechanism in Fig. 1 with substituted connection. The expression for zo in Eq. (60) does not seem to get rid of the twist coordinates for any choice of z expressed in terms of the curvatures of contacting surfaces. This means that the centers of the substitute-connection cannot be obtained using only the curvature values obtained by mere examination of the profiles of the contacting surfaces in a general scenario. But it can be easily shown that substitution using geometry parameters alone is possible for local spherical surfaces in contact, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003591_isc251055.2020.9239095-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003591_isc251055.2020.9239095-Figure2-1.png",
        "caption": "Fig. 2. The power line aerial images with broken strand for each view. The red circles or rectangles highlight the broken strands.",
        "texts": [
            " Downloaded on December 24,2020 at 03:09:25 UTC from IEEE Xplore. Restrictions apply. this section, four aerial views of the power strand to be considered in this work will be introduced first. For each view, an image from the view with broken strand will be presented. Then, the data oversampling and transformation techniques will be described in detail. A. Image Dataset Our dataset consists of the power line aerial images from four typical views. For each view, one image with broken strand is selected to be presented in Fig. 2. Generally, the images captured by UAVs often contain cluttered background with the presence of power towers, cottages, mountains, farmland, rivers, and so on. Fig. 2(a), Fig. 2(b), and Fig. 2(c) refer to the three views sharing one similarity that the broken part is around the end points of a strand and close to the joints of the transmission towers. Among these three views, Fig. 2(a) contains the smallest region of a strand. Also, the background is blurred due to the micro-focus of the camera. Fig. 2(b) contains more infrastructure information as well as the power strand. The background cottage is clear and relatively close to the power lines. This may introduce some difficulty in filtering backgrounds. Fig. 2(c) is the most informative view in our dataset with complicated structure and background; especially that a huge amount of the image contents are contributed by the tower infrastructures where the strand information is relatively minimized. As for Fig. 2(d), it shows a different view type with a long power strand but no other power facilities like tower. the background is also nontrivial with the existence of buildings, trees, grass land, and water area. To prepare the images for ML model training and testing for each of the four views, two issues should be considered about data insufficiency and background denoising. The former is due to the fact that only one raw image is available with the broken strand for each view. The later refers to the complex background",
            " For example, \u2704 \u2702 means the model is trained on view 2 and tested on view 1. As mentioned above, view 1 to 3 have a similar pattern and background, however, become more and more informative. Thus, such three views are used to train and test the ML algorithms. Taken DNN as an example, the performance is shown in Fig. 3. In Fig. 3, DNN trains, and tests on all three image types. The performances on RGB and gray images are still similar to each other. However, the line feature contributes a lot in heterogeneous views which can reach 69.7% in the case \u260e \u2706 \u271d. Refer to Fig. 2, the backgrounds of view 1 to 3 are all different from each other. The line images can well filter out background information and focus on the close shot only which includes the common objects like power line and tower infrastructures. To this end, the ML algorithms are able to apply the model gained while training on one view to a different but related view. Actually, the highest accuracy for DNN is not displayed in Fig. 3 which is 78.1% achieved in the case \u271f \u2720 \u2721. In addition, DNN can only achieve a good performance when it is trained on a complicated view, as view 2 and 3, and tested on a simpler view, as view 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002927_tec.2020.2983187-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002927_tec.2020.2983187-Figure2-1.png",
        "caption": "Fig. 2. Stator of a 12-pole/36-slot claw pole alternator.",
        "texts": [
            " The field winding is driven from the voltage regulator via slip rings and carbon brushes. When the exciting current flows through the field winding, axial magnetic flux is produced. This axial magnetic field flows through the two main pieces of rotor core and each of them becomes a north or south. When the rotor is driven by the belt, the alternating current is induced in the stator winding. The alternating current is later rectified into direct current to charge the battery. B. Types of Stator Winding Interturn Short Circuit Fault of the Same Phase Winding Fig. 2 shows the stator of a 12-pole/36-slot claw pole alternator. In order to maximize the output current and ensure that the output voltage of each phase winding is the same, almost all automotive claw pole alternators are equipped with integer slot waveform windings. In addition, the copper wires wind backwards and forwards in the slots so that the end winding would be evenly distributed at both ends of the stator core, as shown in Fig. 3. The claw pole alternator has a compact structure. In order to improve the power density as much as possible, the stator is required to have a high coil fill factor, that is, the diameter of the copper wire should not be too large"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure6-1.png",
        "caption": "Fig. 6 The mesh map of the part of the gear pump gear teeth in cycle three",
        "texts": [
            " The element effect matrix of H~P e4 and H~P e10 is obtained through interpolating in the solving domain, and the new response matrix H~1 can be obtained through combining H~P e4 with H~P e10 , and the damage coefficient matrix can be expressed as follows: D~ \u00bc \u00bdd11; d12; d13; d14; d15; d16 T \u00bc \u00bd 0:018; 0:3126; 0:0195; 0:0538; 0:0672; 0:0336 T The component which is negative value can be set as 0, and the new damage coefficient matrix can be expressed as follows: D~1 \u00bc \u00bdd12; d15 T \u00bc \u00bd0:075; 0:023 T The precision error is set as n = 10, and the same pro- cedure is carried out again. Cycle 2: the wavelet finite element model of the gear pump gear is shown in Fig. 5. D~2 \u00bc \u00bdd22; d27 T \u00bc \u00bd0:1632; 0:5467 T Cycle 3: the meshing situation is shown in Fig. 6. Intact xi 9.793 23.236 57.294 86.664 112.753 176.475 Cracked xi \u2019 9.361 22.873 56.541 85.357 111.532 175.352 D~3 \u00bc \u00bdd32; d35 T \u00bc \u00bd0:5832; 0:8836 T The zones P e32 and P e35 are identified as cracked element, and the angle is less than the error, and the iteration ends, the precious location, and the identification of the depth is carried out according to the contour line method. The damage parameters are defined as D~3 \u00bc \u00bd0; 0:6326; 0; 0; 0; 0 T for zone P e32, which are put into expression (32), and the changing rate of the natural frequencies of the multi-cracked gear pump gear Dxr xr corre- sponding to D~3 is calculated, and the results are listed as follows: Dx1 x1 \u00bc 0:153 %; Dx2 x2 \u00bc 0:457 %; Dx3 x3 \u00bc 0:338 % The crack identification of the gear pump gear can be carried out according to the variety of the natural frequencies of the gear with multi-cracks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000025_rnc.4734-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000025_rnc.4734-Figure2-1.png",
        "caption": "FIGURE 2 Interception geometry",
        "texts": [
            " The angle of attack \ud835\udefc can be calculated via (1), and other angles can be measured by inertia sensors on the controlled missile. Remark 3. Assume that 4 fins are configurated as \u201c+\u201d structure, as shown in Figure 3, and then, fin deflection can be calculated by \ud835\udeff = (\ud835\udeff1 \u2212 \ud835\udeff3)/2. Because interceptor is an axisymmetric structure, fin deflection \ud835\udeff is equal to \ud835\udeff1 and \ud835\udeff3 numerically. Assuming that both the interceptor and the target are point masses, the planar homing engagement dynamics is depicted as in Figure 2, where M represents the interceptor and T represents the target, and other denotations can be found in the nomenclature. According to Figure 2, some equations can be concluded to describe the interception geometry at the end game of guidance. .r = VT cos (\ud835\udefeT \u2212 \ud835\udf06) \u2212 VM cos (\ud835\udefeM \u2212 \ud835\udf06) . \ud835\udf06 = [VT sin (\ud835\udefeT \u2212 \ud835\udf06) \u2212 VM sin (\ud835\udefeM \u2212 \ud835\udf06)] \u2215r . \ud835\udefeM = aM\u2215VM . \ud835\udefeT = aT\u2215VT (3) Taking the time derivatives of the first and second equations in (3) yields r\u0308 = r . \ud835\udf06 + a\ud835\udc47 \ud835\udc5f \u2212 aMr ?\u0308? = \u22122 .r . \ud835\udf06 r + a\ud835\udc47\ud835\udf06 r \u2212 a\ud835\udc40\ud835\udf06 r , (4) with aTr = aT sin(\ud835\udf06 \u2212 \ud835\udefeT), aMr = aM sin(\ud835\udf06 \u2212 \ud835\udefeM), aT\ud835\udf06 = aT cos(\ud835\udf06 \u2212 \ud835\udefeT), and aM\ud835\udf06 = aM cos(\ud835\udf06 \u2212 \ud835\udefeM). By denoting Vr = .r and V\ud835\udf06 = r . \ud835\udf06, (4) can be rewritten as "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002477_etcm.2016.7750839-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002477_etcm.2016.7750839-Figure1-1.png",
        "caption": "Fig. 1. Basic structure of a Quadcopter. = \u2212 \u2026\u2212 \u2026\u2212 \u2026 \u2026 +\u2026 \u2212\u2026 (1) Where ( , , ) represent the linear positions and (\u2205,, ) represent the roll, pitch and yaw angles respectively. The first reference system ( ) is located on the ground and the second system ( ) is located in the center of mass of the quadcopter.",
        "texts": [
            " The controllers are used for stabilization and trajectory tracking. To analyze the stability, the direct method of Lyapunov is used. This paper is organized as follows. Section 2 presents the kinematic model of the quadcopter. Section 3 shows the design of the controllers. Section 4 presents the results of the simulation tests using the designed controllers. Finally, section 5 presents the conclusions of this work. II. KINEMATIC MODEL OF A QUADCOPTER This section describes the kinematic model of a quadcopter, as show in Fig.1 witch is going to be used in the design of different kind of controllers that will be used and compared in the tracking trajectory of the quadcopter formation. The complete kinematic model, is described in (1): 978-1-5090-1629-7/16/$31.00 \u00a92016 IEEE For the reduced kinematic model [12], a technique called \u201csmall angle approximation\u201d is used, which is a very convenient technique to simplify trigonometric laws and presents an acceptable accuracy when the angle approaches to zero, this model assumes that \u2192 0 and \u2192 0, therefore, \u2245 1, \u2245 1 and \u2245 0, \u2245 0, given the next equation as follows: = \u2212 000 0 1 (2) III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure6-1.png",
        "caption": "Fig. 6 Structure sectioned along the long axis. Cutting plane is left visible so the thickness can be easily seen. Other dimensions: semimajor axis hair base diameter (not shown): 10.4 lm; total hair length: 1100 lm; tissue, iris, and top and center socket thickness: 0.5 lm.",
        "texts": [
            "2) for analysis of the mechanics of the hair\u2013socket assembly under a small force applied to the hair. To ensure greater control during the meshing process in ANSYS, a symmetry plane was employed on the entire model and additionally the socket was also split into parts (bottom cavity, skin for boundary condition, middle, and top bulges of the socket along with the tissue material). The sectioning of the model vertically and further into different parts in Workbench was to optimize distribution of the limited nodes of ANSYS version that was used and to also reduce computational times. Figure 6 shows cross-sectional image of the solid model. Dimensions shown in Fig. 6 are approximated from the actual measurement performed using IMAGEJ software. 4.2 FEA Model Development. The SOLIDWORKS assembly was imported into the ANSYS modeling interface of Workbench. Due to the symmetric nature of geometry and loading conditions, only half of the assembly was used for FEA. A symmetry boundary conditions were employed on the symmetry plane (the cut section shown in Fig. 6) so that all the nodes on this plane have only the in-plane deformation. 4.2.1 Contact Boundary Conditions. Workbench automatically selected bonded contacts between the hair base and socket structure by default, and these contacts were subsequently refined to insure accuracy. Figure 7 shows the contact and target faces. Journal of Biomechanical Engineering AUGUST 2016, Vol. 138 / 081006-3 Downloaded From: http://biomechanical.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jbendy/935362/ on 03/13/2017 Terms of Use: http://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000318_jfm.2013.30-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000318_jfm.2013.30-Figure1-1.png",
        "caption": "FIGURE 1. Two spheres, linked by a rod of periodically changing length.",
        "texts": [
            " At the same time, the major oscillatory forces available in fluid have not been exploited; these are the forces caused by fluid oscillations which are imposed by periodically varying boundary conditions, waves, or turbulence. The ratio of characteristic spatial scales (several microns for a micro-robot versus millimetres, centimetres, or greater scales for flow oscillations) makes it clear that the first problem to study is the behaviour of a micro-robot in a fluid that oscillates as a rigid body. In this paper, we consider the self-propulsion of a two-sphere buoyancydriven dumbbell micro-robot (which we call a BD-robot), see figure 1. The whole micro-robot is neutrally buoyant (in order to avoid sedimentation); one of its spherical beads is positively buoyant and the other is negatively buoyant. We study two versions of BD-robots. In the first one the beads are connected by a rod of prescribed oscillating length, in the second one the beads are linked by an elastic spring. First, we study the case of a rod and, next, we consider the changes that appear after replacing the rod with a spring. A mathematical formulation of the problem leads us to the study of creeping motion with time-periodic forces",
            " Our calculations show that, generally, the BD-robot undergoes both translational and rotational motion. Rectilinear translational self-propulsion with constant velocity represents a special case of this solution. We have calculated the velocity of rectilinear self-propulsion and the ranges of governing parameters that correspond to translational motion. The BD-robot represents a dumbbell configuration, which consists of two homogeneous rigid spherical beads of different radii R\u03bd , \u03bd = 1, 2 connected by a rod of length l, see figure 1. We study two-dimensional motion of a three-dimensional 717 R8-2 dumbbell in Cartesian coordinates (x, y). The centres of the spheres x(\u03bd) are described by x(1) = X + a, x(2) = X + b, R1a+ R2b= 0, (2.1a) a= a\u03c3 = r1l\u03c3 , b= b\u03c3 =\u2212r2l\u03c3 ; r1 \u2261 R2/(R1 + R2), r2 \u2261 R1/(R1 + R2), (2.1b) where X = (X,Y) is the radius-vector of a centre of reaction. The axis of symmetry of a dumbbell is given by the vector l \u2261 x(1) \u2212 x(2), l \u2261 |l|. The unit vectors \u03c3 ,n and the angle \u03d5 are given by \u03c3 \u2261 ( cos\u03d5 sin\u03d5 ) , n\u2261 ( \u2212 sin\u03d5 cos\u03d5 ) , n= \u03c3 \u03d5, \u03c3 =\u2212n\u03d5, \u03c3 \u00b7n= 0, (2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure8-1.png",
        "caption": "Fig. 8 Contact faces on the hair body (left) and target faces on the socket body for the hair bulge/socket iris contact region (right)",
        "texts": [
            " Because the hair base and socket base have a tight fit where they are touching within the restraining socket base, no separation could occur in any plane, so normal stress transferred between the two contacting bodies would be seen. It is known that a second point of contact exists for cricket cercal filiform hairs. Upon application of an air-current with a high enough velocity, a hair will deflect enough from its rest position to contact the edge of the socket. This contact point occurs between the bulge of the hair and the inner edge of the narrowest constriction, which we have defined as the socket\u2019s \u201ciris\u201d in Fig. 8. The contact region is indicated in this figure with a blue line. In our model, the hair bulge was set as the contact body while the socket rim was the target. This contact region was also set to be frictionless. 4.2.2 Mesh Generation. Before importing the assembly, the model was split into eight parts in SOLIDWORKS. Solid tetrahedron elements were used to model the structure. The two main solid elements used for meshing were SOLID186 and SOLID187. Both these elements are higher order 3D elements that have quadratic displacement behavior and are well suited for nonuniform or irregular meshes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001097_s105261881304002x-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001097_s105261881304002x-Figure1-1.png",
        "caption": "Fig. 1. The rotor of a turbo driven pump assembly with seal rings: (i) = 1, 2, \u2026, n is the number of the finite element, [i] = 1, 2, \u2026, and n + 1 is the number of the unit.",
        "texts": [
            "76 SPECIFIC FEATURES OF A MATHEMATICAL MODEL OF A ROTOR WITH SEAL RINGS In this paper, stationary and transient oscillations in high speed rotor systems in which the operating speeds exceed the critical speeds are studied. The object under investigation is the rotor of a turbo driven pump assembly used in fluid propellant rocket engines (see Fig. 1) the rated speed of which exceeds the second critical speed and in which floating rings serve as sealing. A finite element model was constructed for this rotor in which the shaft segments are considered as beam finite elements considering the shear deformation [1] and the blade wheels and the seal rings as rigid disks and elastically attached solid bodies. It considers the following aspects: Gyroscopic moments. The gyroscopic terms \u03c9I are introduced by a skew symmetric matrix that relates the oscillations in planes (xz) and (yz) and \u03c9I = \u03c9 , where \u03c9 is the rotor speed",
            " For this purpose the condition (3) must be satisfied. K* 0 0 K* 0 K* K*\u2013 0 \u2202 \u2202z \u03b43 kz \u2202p0 \u2202z \u239d \u23a0 \u239b \u239e \u2202 R 2\u2202\u03b8 \u03b43 k\u03b8 \u2202p0 \u2202\u03b8 \u239d \u23a0 \u239b \u239e+ \u03bc 0.5\u03c9 \u03b8\u00b7\u2013( )\u2202\u03b4 \u2202\u03b8 .= z 0, p0 p1 p2\u2013 1 \u03b7 1 \u03c7 \u03b8cos\u2013( )+ and z L, p0 0.= = = = \u03b8\u00b7 khi \u03c0LR\u03b7 2\u03b4 1 \u03b7+( )2 \u0394p, dhi 0.005\u03c0\u03bcL 3 R 12\u03b43 Re,= = \u03c9n 2 khi n \u03c92 /\u03c9n 2( ) khi n Mq\u00b7\u00b7 \u03c9G Dext Dint+ +( )q\u00b7 K \u03c9Dn 0.5\u03c9H+ +( )q+ + F, I0\u03d5\u00b7\u00b7 M,= = \u03bdi khi n /mi 1.1\u03c9n\u2248= 278 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 42 No. 4 2013 BALAKH, NIKIFOROV Let us consider the multi mass rotor system of a turbo driven pump assembly (see Fig. 1) in which the parameters of the seal rings are selected according to expression (3) and the table. Changes in the natural frequencies of this system depending on the speed in the vicinity of the most unsafe second critical speed are illustrated by Fig. 2a. The natural frequencies of the system define the intersections of the curves \u03bb(\u03c9) with the ray \u03bb = \u03c9. The upper curve in this diagram corresponds to the frequency curve \u03bb(\u03c9) of the rotor. Since the hydrostatic stiffness values of n rings are smaller than the rotor\u2019s flexural stiffness, n rings add n basic natural frequencies\u2014indicated by the dotted lines\u2014to the rotor system",
            " In the case when the rotor rotates not in floating but fixed seals, the effect of the disappearance of the critical speed owing to the action of the hydrodynamic forces in the clearances also exists. However, according to the calculations, to achieve this effect, a rather high pressure drop is required, higher than that required for floating rings almost by an order of magnitude. THE DYNAMIC DAMPING OF THE ROTOR\u2019S OSCILLATIONS UNDER UNSTEADY STATE CONDITIONS The unsteady state conditions, namely, speedup, of the rotor (Fig. 1) were studied on the basis of equation (2) at a constant angular acceleration. For transient conditions upon the speedup/slowdown, the torque M is usually given in the form of a joint characteristic of the turbine, viz., the torque M1, and of the rotor, viz., the moment of resistance M2 related to the working load [5] (4)M M1 M2\u2013 Mn 2 \u03d5\u00b7 /\u03c9n\u2013( ) Mn \u03d5\u00b7 /\u03c9n( ) 2 ,\u2013= = where Mn and \u03c9n are the rated values of the torque and the speed. The translational coordinates and the angle of rotation can be considered to be independent with small limits on the order of e/r where r is the rotor radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure4-1.png",
        "caption": "Fig. 4: Saturation map and current density map from the nonlinear TH simulation of the IM LR working point at full voltage and frequency: TH non-linear simulation.",
        "texts": [
            " For this purpose, an equivalent sinusoidal winding is considered in the stator side and non-linear TH simulations are performed. The procedure in Fig. 3 is used to derive the stator current in LR operation, imposing s = 1, |us|= \u221a 2Vn and f = fn. The resulting stator current is |is|= 135 A. The current phasor is imposed in a TH non-linear simulation, in which only the fundamental magneto-motive force is imposed, since an equivalent sinusoidal stator winding is considered. Further, several simulations have been done, moving a bit the rotor, in order to mediate the effect of the slot harmonics. In Fig. 4a the flux and saturation map from the simulation field solution is reported. The most saturated machine part is the rotor and, in particular, the upper parts of the teeth. The saturation is mainly due to the leakage flux, instead of the magnetizing one. As a manner of fact, the rotor current, in the LR operation, tends to almost totally shield the stator field, so they are almost in opposition of phase. Only a little part of the stator current is meant to magnetize the machine. This fact cames up also from the procedure in Fig. 3: with s = 1, the d-axis current, that magnetize the machine is isd = 3.3195 A, whereas the rotor current referred to the stator is isq = 135.8256 A. In Fig. 4b, the current density map is shown, in particular concerning the real part, that represents a time instant of the machine operation. The non-uniform current distribution within the rotor bars can be observed. From the field solution, the bar currents distribution along the rotor periphery can be derived and it is reported in Fig. 5a. Both the real and the imaginary parts show a sinusoidal behavior and the current amplitude is the same in all the rotor slots. Thus, it can be concluded that, with a single harmonic field excitation, even in presence of high saturation, the rotor reaction is spatially sinusoidally distributed",
            " They are almost equal each other, thus, with a single harmonic excitation the rotor induced current has to be sinusoidally distributed, with the same number of poles as the magneto-motive force. From Fig. 5b, it can be noticed that the reactive component of the bars impedance is rather important, with a relatively high rotor frequency. Further, the bars reactances are very similar each other. It is like the saturation does not affect that much the leakage inductance of the bars. As a manner of fact, looking at Fig. 4a, some bars are surrounded by heavily saturated iron and others by iron with very low flux density. This could be due also to how the current itself is distributed inside the rotor slots, leading to constant equivalent parameters of each bar. This is an interesting result for the MS analysis of the IM: considering a certain field harmonic, the rotor induced Authorized licensed use limited to: Carleton University. Downloaded on June 30,2020 at 13:07:46 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (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. current by this harmonic is sinusoidally distributed, along the rotor angular coordinate, for any level of saturation and rotor frequency. Thus, the hypothesis made in [14], about the spatial behavior of the rotor reaction is validated. The non-linear TH simulation in Fig. 9a and Fig. 4b represent the benchmark for the procedure described in the following, that have the scope to include the skin effect in the MS analyses of the IM. In the TH analyses, the rotor bars are modeled as plain conductors and no circuit property are imposed. A more complete TH simulation is reported for the final comparison, in which the whole machine is modeled and a cage circuit is defined to guarantee that the sum of the bars currents is zero. A further step for the MS analyses is including the effect of the actual rotor frequency to the rotor current: it can cause a non-uniform distribution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000764_s40032-014-0147-8-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000764_s40032-014-0147-8-Figure1-1.png",
        "caption": "Fig. 1 Circular bearing geometry",
        "texts": [
            " In the literature, no investigation is found addressing the turbulence flow for micropolar lubricant. Hence, the aspect of turbulent lubrication using micropolar lubricant has a great scope for investigation to fulfill the gap. The main focus was to analyze the performances of the system. The modified Reynolds equation developed by prior researchers [16] has been considered for the present analysis, in a similar approach of works carried out by various literatures [14, 15] for couple stress fluid. The circular bearing geometry is shown in Fig. 1. The variation of the film thickness along the circumference of a journal bearing is, h \u00bc h C \u00bc 1:0\u00fe e cos h ; \u00f01\u00de where h is the angular coordinate starting from the line of centers. When the film oil is a non-Newtonian fluid, the Reynolds equation can still be successfully used by means of some modifications for turbulent micropolar lubrication. One possible way of taking into account the influence of the non-Newtonian behavior is the introduction of the micropolar function in the Reynolds equation as described by various researchers [13]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003403_s1068366620040145-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003403_s1068366620040145-Figure1-1.png",
        "caption": "Fig. 1. Universal bench for determining the static stiffness characteristics of tires: (a) general view; (b) kinematic diagram: (1) test wheel; (2) device for installing wheel of given size; (3) position fixers; (4) device for normal (radial) loading of wheel; (5) cart; (6) wheel lateral loading device.",
        "texts": [
            " The second were designed to study the wear rate of the same tires in road conditions. Three tire variants of the size 11/70R22.5 were used in the tests: model no. 1 of domestic production and foreign tire model nos. 2 and 3. 354 To determine the normal and lateral stiffness coefficients, the forces and displacements in the spot of contact between the tire and the solid support were measured for the same-named coordinates. Tests for determining the static characteristics of tires were carried out at the test bench at the NICIAMT tire laboratory (Fig. 1). The design of the bench made it possible to apply forces from 0 to 50 kN (5 tf) to the tire and achieve loading with normal (radial), lateral, tangential (longitudinal), torsional, and angular loads in different combinations. During the experiment, the applied forces and displacements caused by these forces were measured at the spot of contact between the tire and base plate of the bench for the same-named coordinates. The design of the bench also made it possible to determine the geometric parameters of the spot of contact and the specific power indicators of interaction between the wheel and the supporting surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003733_icems50442.2020.9290866-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003733_icems50442.2020.9290866-Figure4-1.png",
        "caption": "Fig. 4. Laminated core shape",
        "texts": [
            " Downloaded on August 13,2021 at 07:27:59 UTC from IEEE Xplore. Restrictions apply. it can be satisfied that full power driving of the proposed motor can be continued for several minutes. The field coil is designed that current density on maximum field current is not exceed than 30 Arms/mm2. Since field coil of the proposed motor is assumed for assisting function of the exciting winding, coil designing is conducted that continuous driving time on maximum field current is lower than that of the exciting windings. In Fig. 4, rotor and stator core shapes are shown. The leakage bridge in rotor core is set as twice width as board thickness. This value is decided from experience to enable press fitting. 2) Results In Fig. 5, reduction effect of the proposed motor, condition of which is set 5 Nm load for whole operation speed 5000 to 20000 r/min, is shown. Since flux range of the proposed motor is correspond to the maximum and minimum field current condition, it can be seen that characteristic of copper loss reduction can be estimated by using analysis results of both conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003002_s00170-020-05276-z-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003002_s00170-020-05276-z-Figure15-1.png",
        "caption": "Fig. 15 Welded combustor casing. a Meridian plane. b External casing. c Diffuser and internal casing",
        "texts": [
            " This means that the simulation results are applicable to engineering practice and prove the feasibility of using a combined ellipsoidal and Gaussian rotating body heat source model to simulate the EBW process. The results also affirm the superiority of the model\u2019s capacity to reflect real welding conditions and improve the accuracy of numerical calculations. According to the EBW heat source simulation and residual stress calculation, the low-cycle life prediction of welded combustor casing was also carried out in this study. A typical combustor casing from a civil aviation engine was used as an example (see Fig. 15). Like the welded plate used in Section 2 above, the thickness of the casing was 2 mm, so the EBW parameters remain unchanged. Given the operational load spectrum during actual flights, including the temperature (see Fig. 16), pressure, and axial force. The welding residual stress was added into the welded joint calculations in the form of a boundary condition, and its specific value was consistent with the calculated results in Section 3.1. TheManson\u2013Coffin [38\u201340] formula shown in Eq. (7) was used to predict the low-cycle fatigue (LCF) life of the welded combustor casing: \u03b5a \u00bc \u03c3 0 f 2Nf\u00f0 \u00deb=E \u00fe \u03b5 0 f 2Nf\u00f0 \u00dec \u00f07\u00de Here, \u03b5a is the strain range; Nf is the structural low-cycle fatigue life; \u03c3 0 f is the fatigue strength coefficient; \u03b5 0 f is the fatigue plasticity coefficient; b is the fatigue strength exponent; c is the fatigue plasticity exponent; and E is the elasticity modulus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002892_ki48306.2020.9039870-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002892_ki48306.2020.9039870-Figure1-1.png",
        "caption": "Fig. 1. Twisting proces of the fishing line actuator [1].",
        "texts": [
            "78-1-7281-4381-1/20/$31.00 \u00a92020 IEEE Keywords\u2014artificial muscle, fishing line actuator, mechanical properties I. INTRODUCTION Twisted and coiled fishing line actuator represents a soft actuator (also called as artificial muscle) which is suitable to realize smooth and soft motion of machines and robots. This type of actuator is fabricated by twisting nylon fibres (fishing line) into helical state and in the final it looks like a preloaded tensile spring (Fig. 1). Cold actuator is stretched due to tensile load and its shortening is realized by heating. Thus, the actuator's operation can be controlled by changing its temperature. Actuator heating is realized by hot air blowing, warm water bypass or by electric Joule heating (Fig. 2). The advantage of this actuator is simple and cheap production, outstanding power density and large deformation. These actuators are referred to as Twisted-Polymeric-Fibre Actuators (TPFA), while the nylon fishing line is most commonly used to produce them"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002020_tmag.2015.2447527-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002020_tmag.2015.2447527-Figure8-1.png",
        "caption": "Fig. 8 The asymmetrical eccentricity optimization of edge PM.",
        "texts": [
            " 0.00 3.75 7.50 11.25 15.00 18.75 22.50 26.25 30.00 -0.5 0.0 0.5 1.0 1.5 90mm 91mm T or qu e (N m ) Rotor Rotation Angle (mech. deg) 0mm 80mm 88mm (a) 0018-9464 (c) 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In order to further reduce the cogging torque, a new technique adopting asymmetrical shape for the first edge PM is proposed, as shown in Fig. 8. Firstly, for a given thickness of the first edge PM, only the outer half pole adopting different eccentric distance is researched while keeping the inner half pole unchanged. Then only the inner half pole adopting different eccentric distance is researched. Finally, the first edge PM adopting asymmetrical shape with different thickness is also researched and compared. Figs. 9-11 show the cogging torque waveforms of different thicknesses. It is found that with the increase of the outer half pole eccentric distance, the cogging torque during the outer end part has the tendency of varying from minus value to plus value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001427_j.oceaneng.2014.06.019-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001427_j.oceaneng.2014.06.019-Figure1-1.png",
        "caption": "Fig. 1. Body-fixed and inertial coordinate systems.",
        "texts": [
            " The proposed controller is a trajectory tracking controller, so it offers optimal time as long as its references (inputs) are the time optimal trajectories, even with uncertainties. 2. Equations of motion Full equations of motion of an underwater vehicle can be found in Fossen (1994). However, in this paper, because we just consider the design of time optimal trajectories for the depth control of a fully actuated UUV, we only need to consider the body-relative heave velocity w, and the earth-relative vehicle depth z (see Fig. 1). We will set all other translational and rotational velocities to zero, and assume that the roll, pitch and yaw angles of the vehicle always are kept at zero for simplicity. As a result, the mathematical model of depth motion of the UUV is as follows: \u00f0m Z _w\u00de _w Zwjwjwjwj \u00bc \u00f0W B\u00de\u00feZprop \u00f01\u00de _z\u00bcw \u00f02\u00de where, m: UUV mass, W: vehicle weight, Z _w: added mass coefficient, B: vehicle buoyancy, Zw|w|: cross-flow drag coefficient, and Zprop: thrust force. Substituting Eq. (2) into (1), we have the nonlinear second order differential equation of depth motion as follows: \u00f0m Z _w\u00de\u20acz Zwjwj _zj_zj \u00bc \u00f0W B\u00de\u00feZprop \u00f03\u00de Setting a\u00bcm Z _w40; b\u00bc Zwjwj40; N\u00bc B W40 (net buoyancy), and u\u00bc Zprop, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000108_ffe.13138-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000108_ffe.13138-Figure5-1.png",
        "caption": "FIGURE 5 Contour plots of (A) contact pressure generated at the interface and (B) von Mises stress generated in the rod for the loading conditions in the calibration data set. The contact had semimajor and semiminor axis dimensions of 0.43 and 0.24 mm, respectively",
        "texts": [
            " Since ceramic materials have a high hardness/yield strength, the response of the ball was assumed to be linear elastic. Table 1 summarizes the material properties of the ball and rod. The radial load was applied normal to the ball\u2010rod contact. Symmetry boundary conditions were appropriately utilized at the ball and rod sections to simulate quarter symmetry. The analysis generated an elliptical contact patch with its major axis along the axis of the rod. The semimajor and semiminor axis dimensions of the contact were found to be 0.43 and 0.24 mm, respectively. Figure 5 shows contour plots of the contact pressure generated at the interface and the von Mises stress generated in the rod. A mesh convergence check conducted on this model has been detailed in Appendix A. acks. (B) The Weibull chart used by Londhe et al. to calculate the L10 To obtain the macroscale stress history \u03a3(x,t) that points at a certain depth in the rod experience under one stress cycle, the six components of stress were extracted along the paths of constant depths beneath the contact as shown in Figure 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure12-1.png",
        "caption": "Fig. 12: First stored saturation map used in the following incremental permeability TH simulation.",
        "texts": [],
        "surrounding_texts": [
            "Let\u2019s come back to the LR operation example, considering now the actual stator uniformly distributed winding. The harmonic fields are rather important for determining the saturation map and the starting torque because of the high stator current. In this work, the strategy adopted to compute the rotor harmonic currents is based on TH simulations, in which only the belt harmonic under investigation is imposed is the field problem. The rotor reaction is then stored considering the amplitude and spatial position of the induced current and its distribution inside each slot. The adopted procedure to get the rotor reaction to the stator belt harmonics is outlined hereafter. 1) Saturation Map Storing: to consider the saturation effect, incremental permeability TH simulations are performed: the stator spatial harmonic is imposed in \u201can already saturated\u201d and linearized problem. A first estimation of the saturation map comes from the procedure Fig. 3, in which the rotor harmonic windings are supplied to completely shield the stator field harmonics, as a first hypothesis. 2) TH Simulations: in this stage the rotor reaction to the stator belt harmonics has to be determined, performing frequency domain simulations. In the TH problem, the field harmonic is reproduced properly imposing the current in the stator slots; in particular, the real part of the current density distribution has to produce a magnetic field in the same direction that the stator harmonic has in the considered time instant in the MS simulation. The imaginary part is imposed just to consider the rotation of the field with respect the rotor. The field solution is shown in Fig. 13 and it is possible to store the induced current in each portion of each rotor slot, to impose them later on the MS simulation. Authorized licensed use limited to: Carleton University. Downloaded on June 30,2020 at 13:07:46 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. 3) Final MS Simulation: the complete field solution is obtained supplying the harmonic windings according to the previous TH simulation. In Fig. 14a, the final MS field solution is reported, considering the rotor reaction to the belt harmonics and the actual bars current distribution. (a) Current density distribution in the final MS LR simulation, considering the rotor reaction to the stator belt harmonics and the actual bars current distribution. (b) Field solution of the TH simulation at standstill. The result is reported for comparison with the MS approach, in particular about the current distribution within the rotor slots. The main results of the simulation are reported in Table II. Fig. 14: Comparison between the MS approach and the complete TH simulation that include the eddy currents formulation in the FE problem. In order to verify the MS result, a non-linear TH simulation has been performed, considering the actual stator winding and imposing the same stator current as in the MS simulation. The checked value is the stator voltage to validate the TH LR analysis. From the TH LR non-linear simulation, the stator i-th phase voltage is determined as: us,i = 2\u03c0\u03c9s \u03bbs,i + z\u0307s is,i (15) where \u03c9s is the stator angular frequency, \u03bbs,i is the phase flux linkage, is,i is the phase current and z\u0307s = Rs + j2\u03c0\u03c9s Lew. The results of the TH simulation shown in Fig. 14b are reported in Table II, the phases voltages respect the rated vale and the torque can be evaluated. For this reference TH simulation, the full model has been used and each rotor bar, modeled as a plain conductor, is part of a single circuit property, that models the cage, to have the bars\u2019 currents sum equal to zero. Further, considering the locked rotor working point, all the stator harmonics induce currents with the same frequency in the rotor side, and this makes the frequency-domain simulation correct to analyze this kind of working condition. The resulting torque from the field solution in Fig. 14a in TMS = 111.6 N m, very similar to the value obtained from the complete TH non-linear simulation, verifying the effectiveness of the proposed strategy. For comparison, a complete current driven TD simulation has been carried out, imposing the same current used in the TH and MS locked rotor analyses. With 90 simulation per period, the steady-state regime is gotten after 16 electrical periods, for a total computational time around 120 min. The TH simulation last 2.61 min and the overall proposed procedure requires 0.24 min to get the solution. Besides the simulation time saving respect the TD approach, the procedure allows to compute the rotor resistance using (6), as function of the current distribution. Further, the total leakage flux accounts for the rotor iron saturation and the rotor bars skin effect."
        ]
    },
    {
        "image_filename": "designv11_30_0000521_icuas.2015.7152412-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000521_icuas.2015.7152412-Figure2-1.png",
        "caption": "Figure 2. An illustration of inflow model",
        "texts": [
            " Inflow model Compared with the open air rotor duct provides additional lift force due to the promoted air suction over the lip, and leads to a flow turning effect due to the changed momentum of the passing free steam. An inflow model was developed which is a modification of basic momentum theory inflow of an open rotor by Eric and Joseph [3]. The model takes thrust augmentation and flow turning effect into account. An illustration of velocity and force vectors of the inflow model is displayed in Figure 2. The velocity vectors upstream of the rotor, at the disk of the rotor, and far downstream of the rotor are given by 0 0 0 R 0 0 0 0 V cos i sin j V cos i sin j V cos i sin j R R V V V V V V               (1) Where \u03b1 R and \u03b1\u221e are the affected angles and are given by 2 2 R Rk k                           (2) We assume that the total force vector is consists of a part of vertical thrust and another part of horizontal momentum drag. T=Dmi Tj (3) And the thrust component includes rotor thrust and duct thrust"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000228_j.actaastro.2019.12.021-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000228_j.actaastro.2019.12.021-Figure9-1.png",
        "caption": "Fig. 9. The position of the TER tracker mounting on the spacecraft.",
        "texts": [
            " 8, the condition for the satellite to enter the earth shadow350 region is351 \u03c8 \u2265 90\u25e6 + \u03be, (54) where \u03c8 is the angle between w and Ss; w is the unit direction vector of352 the spacecraft from point Os, and w = RT os [ 0 0 \u22121 ]T ; Ss is the unit353 direction vector of the sun from point Os; \u03be is the shadow angle of the earth,354 \u03be = arcsin (\u221a R2 \u2212R2 e / R ) ; R = a (1\u2212 e cos (E)) is the distance from point355 Os to Ob.356 Fig. 8. The earth\u2019s shadow region. 21 5.3. The apparent trajectory of the sun with respect to the TER tracker357 The TER tracker is mounted on the spacecraft, as shown in Fig. 9. The358 direction vector of the sun in O \u2212XY Z can be computed by359 Sd = Rz (\u221290\u25e6)Rx (\u221290\u25e6)So (55) 22 6. Simulation analysis366 6.1. Parameter setting367 Prior to the simulation analysis, it is necessary to set relevant parameters.368 The spacecraft uses a hexahedron satellite with dimensions of 2.02 m\u00d72.00369 m\u00d72.22 m, and a mass of 985 kg [39]. Table 1 lists the orbital parameters370 of the satellite [40]. The structural parameters of the CMR and ECD of371 the SSPCI are listed in Table 2, and those of the TER tracker are shown372 in Table 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003116_calcon49167.2020.9106511-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003116_calcon49167.2020.9106511-Figure2-1.png",
        "caption": "Fig. 2. The components of ball bearings, load zone,force applied and the distribution of loads [7]",
        "texts": [
            " Parallelly applied force to the shaft is named axial force and sheer applied force to the shaft is understood as radial force. So, for up the lifetime of bearings the foremost small print to appear once of this instrumentation are correct alignment, correct placement and enough lubrication [7]. ISBN: 978-1-7281-4283-8 502 PART NO.: CFP20O01-ART Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 05,2020 at 01:16:31 UTC from IEEE Xplore. Restrictions apply. An example of a typical deep groove bearing is given in Fig. 1. In Fig. 2. we can see that a needle bearing consisting of an outer race, an inner race, balls, a shaft and a cage holding the balls. In the figure, it is often ascertained that the zone of load and the distribution of load are given along the direction of applied force. Virtually, the outer race is the stationary part and the inner race and the balls are the rotating part. As the outer race is exposed to load, cracks and pits may happen. The rotating parts may contain other defects due to constant rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000991_s00170-015-6812-0-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000991_s00170-015-6812-0-Figure2-1.png",
        "caption": "Fig. 2 The experimental setup for laser remelting process",
        "texts": [
            " The substrate material was quenched at 1030 \u00b0C and then was tempered at 540 \u00b0C, which is in the same state with the die casting die. The microstructure of the substrate is composed of tempered martensite and granular carbide. 3.2 Sample preparation The specimen used for experiments was a plate with size of 100\u00d750\u00d75 mm. The specimen was mechanically ground and polished, using progressively finer grades of silicon carbideimpregnated emery paper prior to laser remelting process. The schematic of the laser and the experimental setup used in this work are shown in Fig. 2. A pulsed solid-state Nd:YAG laser system (fiber-optic Welder WF300, Han\u2019s Laser, China) with a maximum mean laser power of 300 W and 1.064 \u03bcm wavelength was used as the laser source, using a circularpattern Gaussian beam. Important parameters of remelting for the laser beam are electric current, pulse duration, frequency, defocus distance, scanning speed, and focal length. In this paper, the focal length of the focusing lens is 70 mm. The laser head is fixed vertically on a cylinder and can move up and down in the z-axis direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002569_978-3-319-51691-2_16-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002569_978-3-319-51691-2_16-Figure1-1.png",
        "caption": "Fig. 1. Screenshot of the self-designed experiment setup, (a) standard equipment setup, (b) PCI based AE system for data acquisition.",
        "texts": [
            " Finally, the selected discriminant feature vector is tested for a low-speed rolling element bearing fault diagnosis application. The remaining parts of this paper are designed as follows. Section 2 explains the experiment setup and AE signal acquisition technique. Section 3 describes the proposed fault diagnosis scheme with hybrid feature selection methodology. Experimental results are given in Sect. 4, and concluding remarks are in Sect. 5. The standard scheme for measuring the AE signal is introduced in Fig. 1. We employ some of the most widely used sensors and equipment in the real industries. To capture intrinsic information about defect-bearing and bearing with no defect (BND) conditions, the study records AE signals at 250 kHz sampling rate using a PCI-2 system that is connected with a wide-band AE sensor (WS \u03b1 is from Acoustics Corporation of Physical [5]). The effectiveness of experiment setup and datasets can be studied further in [5, 14]. In this study, AE signals are collected for formulating four experimental datasets of different crack sizes (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure21-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure21-1.png",
        "caption": "Fig. 21. Some special configurations without considering interference.",
        "texts": [
            " 17 ), when the axes of R joints J and J coincide, the axes of R joints J and J coincide, and the axes of R joints J and R2 R4 R6 R8 R1 J R5 coincide during the movement, the mechanism makes a planar motion, which is called plane-motion Mode V. Its typical configurations are shown in Fig. 18 . Due to the characteristics of antiparallelogram unit ( Fig. 19 ), it is a bifurcation point at \u03b8AA = 0 or \u03b8AA = \u03c0 . At a bifurcation point, an antiparallelogram could turn into a parallelogram. Therefore, there are some more special configurations as shown in Fig. 20 (no configurations interfere with each other by the layering). In addition, when interference is not considered, some special configurations are shown in Fig. 21 . Some configurations of plane-symmetric Mode I, plane-symmetric Mode II, plane-motion Mode III, plane-motion Mode IV, and plane-motion Mode V can be transited to each other, and they have a certain relationship as shown in Fig. 22 . It can be obtained that the plane-symmetric Mode I includes configurations (1)-(9); the plane-symmetric Mode II includes (6), (8), (10)\u2013(17); plane-motion Mode III includes (7), (27), (31), (35), (37); plane-motion Mode IV includes (5), (9), (26), (28)\u2013 (30), (32)\u2013(34), (36); and plane-motion Mode V includes (14)\u2013(25)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002632_tencon.2016.7848445-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002632_tencon.2016.7848445-Figure7-1.png",
        "caption": "Fig. 7: Fused deposition modelling process[3]",
        "texts": [
            " SLS printed parts are the stronger than the SL parts as the materials are fused at high temperature and treated after the printing process. Parts that require good mechanical strength like turbine blades, columns can be printed as shown in Fig. 6. SLS provide good mechanical strength as it uses metal powders like Aluminum, Titanium, alloys etc. 2328 2016 IEEE Region 10 Conference (TENCON) \u2014 Proceedings of the International Conference Fused deposition modelling[3][7] is simple technique which utilizes thermoplastic filaments as the source of material. (shown in Fig. 7). Same as all the 3D printers, FDM machine also requires the STL files as the input. The machines has a printer head wherein the thermoplastic filament is held at a nozzle at temperature higher than the melting point of the material at the tip of the nozzle. According to the STL code, the printer head moves in 2D layer over layer melting the material and sticking it. In this machine the platform is stationary, the nozzle head moves in all 3 axes to build the parts. The process is relatively slower compared with the previous two and the mechanical strength is also not an advantage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure8-1.png",
        "caption": "Fig. 8. Composite tooth flank defined by a family of involutoid segments.",
        "texts": [
            " In the two cases presented, the transverse curve is identical. The family evolutoids define a surface, the base surface (family of evolutoids 3 ). Images of the base surfaces for CW and CCW rotation are presented in Fig. 7 . Base surfaces can be entirely on either side of the pitch surface (for different directions of rotation) or intersect the pitch surface as illustrated in Fig. 7 . Associated with each evolutoid is its involutoid curve. A segment of each involutoid can be used to define conjugate tooth flanks. Fig. 8 shows the tooth flanks relative to the base 3 One option is to identify the family of evolutoids as an evolutode . Continuing, the active tooth flank or family of involutoids becomes an involutode . Also, we will have a CW-evolutode and CW-involutode together with a CCW-evolutode and CCW-involutode. surfaces of Fig. 7 . Be careful to note that the active tooth flank is viewed from the toe end for CW pinion rotation whereas the active tooth flank is viewed from the heel end for CCW rotation. It was mentioned earlier that planar involute and evolute curves are partner curves where one curve can be defined in terms of the its partner curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003367_s00170-020-05832-7-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003367_s00170-020-05832-7-Figure1-1.png",
        "caption": "Fig. 1 Jaeger ribbon moving heat source model",
        "texts": [
            " A large number of abrasive particles pass through the grinding zone rapidly, making heat continuously conducted to the cycloid gear tooth surface. Therefore, directly cause cycloid gear tooth surface in grinding area damage. Due to the complex curve of cycloid gear tooth surface, the moving heat source above will cause much higher error in temperature field simulation of cycloid gear form grinding. Since the size of the cycloid gear tooth surface is much larger than that of the grinding area, the grinding area can be seen as a rectangular strip moving heat source along a plane at the speedVw. As shown in Fig. 1, a strip-shaped heat source with width of 2L, the heat source continuously generates heat at a heat intensity of qmcal (1cal= 2.54J) per unit area per second and moves along the X direction at a speed Vw on a plane body. When the rectangular strip moving heat source reaches 0 point (t= 0), keep a stable state. Before time t, the strip heat source is at the position of point (x\u2032\u2212 vt, 0, 0), at this time, the strip-shaped heat source qmdx \u2032dt with the width dx\u2032 makes the temperature of the point p(0, 0, z) increased to d\u03b8 x;z\u00f0 \u00de \u00bc qmdx 0 dt 2\u03c0\u03bbt exp \u2212 x\u2212x0 \u00fe vt 2 \u00fe z2 4at ( ) \u00f01 1\u00de At steady state (t = 0), the temperature of the point p(0, 0, z) is increased to \u03b8 x;z\u00f0 \u00de \u00bc qm \u03c0\u03bb \u222bl\u2010le \u2010 x\u2010x 0\u00f0 \u00dev 2a K0 \u2010 x\u2010x 0 \u00fe vt 2 \u00fe z2 2a ( ) dx 0 \u00f01 2\u00de Let u = (x \u2212 x')v/2a, X = vt/2a, Z = vz/2a, and L = vl/2a; then \u03b8 x;z\u00f0 \u00de \u00bc 2qma \u03c0\u03bbv \u222bX\u00feL X \u2010L e\u2010uK0 Z2 \u00fe u2 1=2 du \u00f01 3\u00de In the formula, K0 is a zero-order second-class modified Bessel function"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003623_ecce44975.2020.9235618-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003623_ecce44975.2020.9235618-Figure7-1.png",
        "caption": "Fig. 7 Mode (0,1) at 4688 Hz.",
        "texts": [
            " Each natural frequency has its vibration mode in the radial and axial direction. The 18-12 pole SRM has vibration mode (0,1) at 4688 Hz where 0 and 1 indicate vibration mode in radial and axial direction, respectively. In addition, there are mode (2,1) at 2714 Hz, mode (3,1) at 3632 Hz, mode (4,1) at 5106 Hz, mode (5,1) at 6635 Hz, and mode (6,1) at 7936 Hz. The problem in the modal analysis is it is unknown which mode is important. In the experiment, mode (0,1) is found to be the most critical. Fig. 7 (a) and (b) show the pictures of the mode (0,1) at natural frequency of 4688 Hz. These pictures are originally provided as a video to show the movement. The only maximum and minimum snapshots are captured as pictures in Fig. 7. The color in the model indicates the amplitude of deformation. The blue color indicates no deformation, while red color indicates the maximum deformation. In Fig. 7(a), the housing, the stator, and the coil are expanding and shrinking, which is vibration mode 0 in the radial direction. On the other hand, in Fig. 7(b), there is a stretching and shrinking movement in the axial direction, which is vibration mode 1. Fig. 8 shows the ANSYS Mechanical result of deformation of the 18-12 pole SRM model when the motor is excited by the square current. The analysis has been carried out by using ANSYS Harmonics Response coupled with ANSYS Maxwell. It is mentioned in [21] and [24] that most of peak noises of 18-12 pole SRM operated at 1000 rpm occurred at the frequency of 3rd current harmonics and its multiples. Since the fundamental current frequency is 200 Hz, the frequencies with peak noise are 600Hz, 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001860_robio.2015.7419705-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001860_robio.2015.7419705-Figure3-1.png",
        "caption": "Fig. 3: Pursuit strategy. P and T are the current positions of the pursuer and the target, respectively.",
        "texts": [],
        "surrounding_texts": [
            "Biological systems exhibit varying delays, and as such, they inspire a change in strategy as a mechanism to deal with interception in the presence of delay. If the assumption is made that insects, in particular a housefly as studied in [1], has evolved to compensate for its inherent sensorimotor delay, then overall transfer function in (4) should, in effect, be a pure gain. In order to convert (4) into a pure gain, we require: k k\u2032 = \u03b3 or, equivalently, k\u2032 k = \u03c4 (6) That is, the ratio of the derivative and proportional gains of the perception module must be equal to the time constant of the motor system. In a study of housefly pursuit behavior [1], the proportional and derivative gains were determined to be (k = 20s-1 and k\u2032 = 0.7s-2), by correlating measurements of \u03c9 with \u03b8e and \u03c9 with \u03b8\u0307e, respectively. In that study, the time constant of the control system was estimated by assuming it to be a pure delay, and determined the delay by locating the peak of the cross-correlation function between \u03c9 and \u03b8e. This yielded an estimate of 30ms for the sensorimotor delay. From our analysis, we can predict that, for their estimated values of k and k\u2032, the time constant (\u03c4 ) of the motor system should be 35ms, from (6), if the perception module is optimised to compensate for the sensorimotor delay. This predicted value is very close to the measured value (30ms), which provides intriguing evidence for the first time that, at least in the housefly, the perception system is tuned to compensate for the time delay introduced by the sensorimotor system. We will also show that the control law in (1) can be utilised not only for pursuit but also constant bearing interception. Greater detail is provided for the pursuit strategy in Section IIIB, and for the interception strategy in Section IIIC."
        ]
    },
    {
        "image_filename": "designv11_30_0003613_s00773-020-00779-6-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003613_s00773-020-00779-6-Figure3-1.png",
        "caption": "Fig. 3 Coordinate systems of 6-DOF ship motion",
        "texts": [
            " A, B, C and U are system matrix, control matrix, measurement matrix and input matrix, respectively. A classical rudder and fin hybrid control system and the coefficient matrices can be referred in [15] If the output is Y = [ ] T , the transfer function of the model G(s) is shown as where the s is a Laplacian operator. G11(s) , G12(s) , G21(s) and G22(s) are The variables describing the position and orientation of ships are usually expressed in the space fixed coordinate system. The translational velocities and angular velocities are denoted in the body fixed coordinate system in Fig.\u00a03. O-X0Y0Z0 is space fixed coordinate. o-xyz is body fixed coordinate system, where the origin is placed at midship. The designed procedure of multi-input multi-output (MIMO) system is more complicated comparing with single input single output (SISO) system [13, 14]. With the development of computed technology, it can solve complex calculation. Whereas it may lead to design higher order controllers, which can cause the higher the order of controller, (2) X\u0307 = AX + BU Y = CX X = [ v p \ud835\udefe \ud835\udf19 \ud835\udf13 ]T U = [\ud835\udf0e \ud835\udeff]T (3) A = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u22120"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure8-1.png",
        "caption": "Figure 8. Wheel/rail critical contact at point M.",
        "texts": [
            " Clearly, the contact position changes irregularly, however, the distance between the contact point and the centerline of the rail\u2019s cross-section is positive. This means that at different relative positions, contact only occurs between the wheel and the inner side of the rail. Methods to calculate the wheel/rail static contact stress Because the wheel and rail surfaces are composed of different curved surfaces, the Hertz contact theory cannot be directly used to calculate the contact stress. Take one contact point M(xM, yM) for instance, its coordinates are the contact spot\u2019s center coordinates when actual deformation occurs, see Figure 8. It is assumed that there is a profile dividing point N near the point M, moreover, point N is in the contact spot when deformation occurs. On the basis of critical contact, the wheel profile moves downwards for a distance of y, and then the actual deformation at point M is replaced by geometry interference of the wheel/rail profiles, shown in Figure 9. By solving profile equations, the boundary points M1(x,y) and M2(x,y) of the interference area can be obtained, and the equation of a straight line between the two points can be written as y \u00bc ux\u00fe c"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-5-1.png",
        "caption": "Figure 6-5. Quadcopter coordinate frames",
        "texts": [
            " Center of pressure use it for vehicles or objects with high aerodynamic variation (projectiles or aircraft in turbulent conditions, for example). Base point use it with anchored bodies, the base of a robotic arm fixed to the ground, or to a heavy load, for example. 246 Propulsion, propeller, motor, or engine frames are reference frames that describe thrust and torque of each propeller or motor. Note that every vehicle has a propulsion frame. In the case of wheeled robots, it is called a wheel frame. All the frames are related in the following way; see Figure\u00a06-5. The propulsion frame indicates how the thrust and torque of each propeller will affect the center of gravity, where their equivalent effects will be transferred. Once it has transferred the equivalent dynamic effects of all motors, the body frame will be cinematically related to the base frame, through the entire movement of the vehicle. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 247 2. Propulsion matrix: Also known as an allocation matrix, mixer, or matrix relation between speeds (the speeds of the actuators with respect to those of the analysis center of the vehicle)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002571_2168-9806.1000135-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002571_2168-9806.1000135-Figure4-1.png",
        "caption": "Figure 4: Contact forces on crawler shoe, i.",
        "texts": [
            " However, that study did not present the solution results for the external forces, which are used as an input in the determination of crawler shoe kinematic (linear and angular acceleration, velocity and displacement) quantities. Volume 4 \u2022 Issue 2 \u2022 1000135 J Powder Metall Min ISSN: 2168-9806 JPMM, an open access journal The gravity force and distributed load on each crawler shoes are constants and are entered directly into the MSC ADAMS as user input. The contact force however acting on each crawler shoes is a time varying force and is calculated using the inbuilt contact force algorithm in MSC ADAMS. The contact forces consist of normal force, frictional force and frictional torque. Figure 4 shows the contact forces (Fn-normal, Ft-tangential and T-torque) developing between track shoe i and ground [11,17]. These forces will act on the crawler shoe shoe bottom surface at a point I [18] as shown in Figure 4. The normal force vector i T N N.x N.y N.zF =[F F F ] acting at point I for the crawler shoe i, shown in Figure 4, is calculated using the impact function model in MSC ADAMS. In this model, when two solid bodies come in contact with each other a nonlinear spring damper system is introduced to determine the normal force [18,19,20]. max * ( ,0,0, ,1) 0 0 0  \u2212 >= \u2264  e i N kx c x Step x d if x if x F (2) The coulomb friction model in Adams is used for the calculation of tangential friction force vector . . . =   Ti T T x T y T zF F FF shown in Figure 4. Based on this model, the frictional force acting at point I can be expressed as equation (3) [11,18,19]. ( )\u00b5=i i T s NF V F (3) ( )\u00b5 sV = friction coefficient defined as a function of slip velocity vector Vs at contact point I [18,19]. The contact parameters listed in Table 5 are used in the study for calculating normal and tangential forces. The contact force on crawler shoe i is the vector sum of normal and frictional forces given by equation (4). = +i i i N TF F F (4) The friction torque iT about the contact normal axis shown in Figure 4 impedes any relative rotation of shoe i with respect to the ground [19]. This torque is proportional to the friction force i TF [19]. 2 3 =i i TRT F (5) R = radius of the contact area. The time varying contact forces are presented for two types of propelling motion constraints imposed on the crawler track [12]. In the first motion type, the crawler track is given only translation motion, and in the second type, the crawler track is given both translation and rotation motion. In addition, the crawler shoe joint constraint forces and torques and total deformation of the oil sand terrain are presented Volume 4 \u2022 Issue 2 \u2022 1000135 J Powder Metall Min ISSN: 2168-9806 JPMM, an open access journal as part of the solution to the equations of motion",
            " This is because normal force vector is nearly perpendicular to x-y plane and tangential force vector lies on the x-y plane. However, for Part 8 similar to normal force developing large component in the x-direction between 0 and 5.3 s (Figure 6a), the tangential force vector also develop large component along the z-direction during the same time period as shown in Figure 8c with a maximum mean value of 11,887 N. This is due to normal force vector not being nearly perpendicular to x-y plane (contact force diagram in Figure 4) when Part 8 moves from oil sand unit 7 to 12 between 0 and 5.3 s. This increase in the tangential force in the z direction leads to a decrease in the frictional force in x direction for Part 8 as shown in Figure 8a. The maximum magnitude of frictional force along the longitudinal (x), lateral (y) and vertical (z) directions are given by 2.2 \u00d7 105 N, 1.73 \u00d7 105 N and 7.82 \u00d7 104 N, respectively. The variation of frictional torque about the contact normal axis Volume 4 \u2022 Issue 2 \u2022 1000135 J Powder Metall Min ISSN: 2168-9806 JPMM, an open access journal Volume 4 \u2022 Issue 2 \u2022 1000135 J Powder Metall Min ISSN: 2168-9806 JPMM, an open access journal with respect to time is shown in Figure 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001252_transfun.e98.a.626-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001252_transfun.e98.a.626-Figure2-1.png",
        "caption": "Fig. 2 Example of directed graph. (Left) Given directed graph. (Right) Directed graph in which redundant edges are eliminated.",
        "texts": [
            " Here, a modeling method of directed graphs such as Fig. 1(b) is focused as one of the methods for decreasing the computation time. This is because in directed graphs appeared in this paper, redundant edges are in general included. By eliminating redundant edges before solving the MIQP problem, decreasing the computation time will be achieved. In this section, a method for eliminating redundant edges is explained. First, consider the case of single-vehicle systems. As an example, consider the directed graph in Fig. 2 (Left). If the current state is given as 4, then we can obtain the directed graph in Fig. 2 (Right). Comparing these, we see that the number of edges decreases. Thus, based on the current state, edges can be eliminated. This method has been proposed in [16], but in [16], some paths (e.g., the path 5 \u2192 5 \u2192 4 in Fig. 2 (Left)) are not eliminated. In the proposed method, by using not only the current state but also the terminal state, edges are eliminated. In this sense, the method in this paper is a more sophisticated version. We remark that the terminal state is not uniquely determined, and multiple candidates exist. Next, consider the case of multi-vehicle systems. As an example, consider the directed graph in Fig. 3. In Problem 4, we suppose that r1 and r2 are given as r1 = 4 and r2 = 5, respectively. In other words, rendezvous of two vehicles must be achieved at either t4 or t5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000816_melcon.2014.6820563-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000816_melcon.2014.6820563-Figure6-1.png",
        "caption": "Fig. 6. Two-mass system driven by a DC servomotor.",
        "texts": [
            " For this purpose, a multi-mass model of two-arm and four-arm systems is developed for analyzing the vibrations that may appear in the upper extremities and the stability of an athlete\u2019s body. The next part shows an application for the two-mass model on a DC servomotor. The electrical drive system will be modeled by two wheels connected together by an elastic link. In this section, we study a simplified two-mass system made of two wheels connected together by an elastic link. The first wheel which represents the motor shaft is driven by a DC servomotor while the second one represents the mechanical load. The system is represented in Fig. 6. U is the voltage source, I the armature current, Rm the armature resistance, Km the servomotor gain, \u03c60 the field flux, J1 the inertia of wheel 1, \u03c41 the time constant of wheel 1, \u03b8 the angular position of wheel 1, dtd\u03b8=\u21261 the velocity of wheel 1, K the rigidity factor, J2 the inertia of wheel 2, \u03c42 the time constant of wheel 2, \u03c6 the angular position of wheel 2, dtd\u03d5=\u2126 2 the velocity of wheel 2, f1 and f2 viscous friction coefficients, and Cload is the load torque considered as a perturbation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002927_tec.2020.2983187-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002927_tec.2020.2983187-Figure1-1.png",
        "caption": "Fig. 1. Exploded view of a claw pole alternator.",
        "texts": [
            " The outline of this paper is as follows: Section II introduces the structure and the stator winding interturn short circuit fault of the claw pole alternator; Section III illustrates the circuit simulation model and current simulation result under interturn short circuit fault; Section IV analyzes the characteristics of magnetic force and noise under interturn short circuit fault; Section V summarizes the conclusions of this paper. II. ILLUSTRATION OF STATOR WINDING INTERTURN SHORT CIRCUIT FAULT OF CLAW POLE ALTERNATOR A. Structure and Operating Principle of Claw Pole Alternator Fig. 1 shows the exploded view of a 12-pole/36-slot electric excitation claw pole alternator, which is universally used as an automotive generator currently. It is a wound-field three-phase synchronous generator containing an internal three-phase diode rectifier and a voltage regulator. The stator is clamped between two brackets. The rotor consists of a pair of claw poles and a ring-shaped field winding. The field winding is driven from the voltage regulator via slip rings and carbon brushes. When the exciting current flows through the field winding, axial magnetic flux is produced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000041_s11668-019-00763-2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000041_s11668-019-00763-2-Figure3-1.png",
        "caption": "Fig. 3 Finite element models of lower arms in ANSYS: a MF285 and b MF399 tractors",
        "texts": [
            " Then, the CAD simulation files in order to build the FEM were imported into ANSYS V15 Software. Mechanical properties of steel (St 37) which was used to build the lower arms are presented in Table 1. Static Analysis The finite element models were meshed using SOLID187 that is a three-dimensional solid element and has 8 nodes with three degrees of freedom for each node. The models consisted of 614 elements and 16,318 nodes for the lower arm of MF399 tractor and 530 elements and 99,089 nodes for the lower arm of MF285 tractor (Fig. 3). Meshing of models was performed by the free method. The boundary and displacement conditions were applied to arm junction points to the tractors. The values of loads applied to the lower arms of MF399 and MF285 tractors were obtained from field testing under different working conditions. The tensile force testing was performed in order to determine maximum value of forces on the lower arms and its direction while working with chisel, furrower equipment and drill planter. The lower arms supply draft forces which are applied in parallel with the forward direction of tractor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002174_amc.2016.7496397-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002174_amc.2016.7496397-Figure1-1.png",
        "caption": "Fig. 1. Prototype of the assistive robot.",
        "texts": [
            " Our key idea is dividing Research supported in part by Grant-in-Aid for Scientific Research C (25350693) from Japan Society for the Promotion of Science (JSPS). 978-1-4799-8464-0/16/$31.00 \u00a92016 IEEE a body motion into a voluntary movement and a postural adjustment by a body movement vector. Our system applies damping control to the body movement vector direction for reducing the load in a voluntary movement and position control to the vertical direction of the body movement vector for assisting a postural adjustment. II. STANDING ASSISTANCE SYSTEM Fig. 1(a) shows our proposed assistance robot. The system consists of a support pad with three DOFs and a walker. The support pad is activated by our new assistance manipulator, which has four parallel linkages [7]. The patient leans on the support pad and grasps the armrest while standing with assistance (see Fig. 1(a)). In general, fear of falling forward during the standing motion reduces elderly patients\u2019 standing ability [9]. With the proposed scheme, patients can easily maintain their posture during a standing motion without the fear of falling forward. Fig. 1(b) shows the prototype of the proposed robot. The prototype is able to lift patients up to 180-cm tall and weighing up to 150 kg. Furthermore, because of its actuated wheels, the prototype can assist patients in walking. To measure a patient\u2019s posture, the prototype has a force sensor and a laser range finder in its body (see Fig. 1(b)). Our physical activity estimate scheme, which is explained in the next paragraph, requires realtime data regarding its assistance force and the patient\u2019s posture. To measure its assistance force, our support pad has two force sensors on its body that measure padF and armrestF (see Fig. 1(a)). To measure the patient\u2019s posture, we use a laser range finder; thus, special markers do not need to be stuck onto the patient for a motion capture systems. 1) Overview of the Proposed Estimate Scheme: In the linkage model of a human body [10, 11], joint traction is used as an index of a patient\u2019s load. However, this index does not consider the posture of the patient, and in some cases, this index diverges from the experience of nursing specialists, especially when the patient is in a half-sitting posture",
            "  The second, we derive the maximum traction (The knee joint: max k , the waist joint: max k ) which muscles can generate with the posture at this time.  Comparing the two derived tractions, we evaluate the physical activity of the patient, i , which demonstrates how much the patient is required their own physical strength as compared with their maximum power (1). i is the identification character (for example, in the case of the knee joint, i is k): 2) Derivation of the Required Traction: The assistance system is designed in such a way that patients lean on a pad and grasp an armrest while standing with our assistance (see Fig. 1(a)), which means that our system uses the pad to apply force to the patient\u2019s chest and the armrest to apply force to their forearm. These forces move vertically (at the pad) and horizontally (at the armrest). Considering these conditions, we use a linkage model that approximates the human body with our assistance device (see Fig. 3). This model consists of six linkages. The armrest applies the assistance force ( armrestf ) to the center position of Link 1 and the support pad applies the force ( padf ) to the center position of Link 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure4-1.png",
        "caption": "Fig. 4. Substitute-connection between the two surfaces with the centers at K and Ko.",
        "texts": [
            " Let v be themagnitude of v and an be the scalar component of acceleration along the oriented n-line n. Then the signed radius of the virtual sphere, denoted by \u03c1, is given by: 1 The \u03c1 \u00bc v2 an : \u00f01\u00de Now, let us suppose that we have such a pair of points K and Ko. Point K can be instantaneously constrained tomove on the virtual sphere by a rigid link-lwhose end points are K and Ko. A 2-d.o.f. kinematic joint is required at the point Ko to constrain the motion of point K on to the spherical surface. This can be achieved by a U-joint whose axes are orthogonal toKKo\u2194 (Fig. 4). Next, the relative motion of the SFm body w.r.t. the point K contained in it has 3 rotational d.o.f. and a S-joint at K will satisfy this requirement. This substitute-connection does not give rise to any redundant kinematic quantities. Consider this substitute-connection as a part of a spatial mechanism. Let vK=Ko and aK=Ko be the velocity and acceleration difference respectively between points K and Ko.\u03c9 l and \u03bc l be the angular velocity and acceleration of the constraining link-l respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000914_jsen.2013.2264284-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000914_jsen.2013.2264284-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the micro-channel.",
        "texts": [
            " The stainless steel master mold was molded by using the computer numerical control (CNC) technique. Then, the PDMS colloid was prepared by mixing main agent of the PDMS and curing agent at 1:10 weight ratio. The PDMS colloid was injected in the master mold and all air bladders were removed by the vacuum ball. The master mold placed in oven to solidify the PDMS, and the PDMS microchannel was lift-off from the master mold. Finally, the inlet and outlet of the microchannel were drilled with 3 mm diameter holes. The microchannel structure and dimension were shown in Figure 1. And the top and cross-section diagrams of microfluidic device were shown in Figure 2. According to previous study in our research group [15], the polypyrrole (PPy) of the conductive polymer was deposited on silver paste by electrodeposited method. The prepared processes of the PPy films were described as follows: (1) The KH2PO4 and K2HPO4 were mixed as phosphate buffer solution. (2) Allocating the electrolytes of the potassium chloride (KCl) and acetonitrile were 0.1M and 1M, respectively. (3) Allocating the concentration of the pyrrole was 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000873_s11071-013-1230-z-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000873_s11071-013-1230-z-Figure1-1.png",
        "caption": "Fig. 1 A simple illustration of the model-scaled helicopter: reference frames, rotor thrusts, and flapping angles",
        "texts": [
            " The problem, including mathematical modeling and control objective, is formulated in Sect. 2. Detailed controller design process is described in Sect. 3. Main theoretical results are summarized in Sect. 4. A brief introduction of the helicopter testbed is presented in Sect. 5. To demonstrate the effectiveness of the proposed controller, results of several practical flight tests are displayed in Sect. 6. The paper is concluded in the final section. 2 Problem formulation 2.1 Mathematical modeling A simple structure of the model-scaled helicopter is illustrated in Fig. 1. Earth reference frame (ERF) and fuselage reference frame (FRF) are defined conventionally [8]. 1. The ERF is fixed to earth, with its origin O locating at a fix point. The x axis points to the north and the z axis points upright vertically. The y axis is confirmed by right-hand rule. 2. The FRF is fixed to fuselage, with its origin Ob located at the center of gravity (CG) of fuselage. The xb axis points to fuselage head. The zb axis is perpendicular to xb axis and points upright. The yb axis is confirmed by right-hand rule"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003631_ever48776.2020.9242986-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003631_ever48776.2020.9242986-Figure1-1.png",
        "caption": "Fig. 1. Cross-sections of 2-pole high-speed motors with toroidal windings.",
        "texts": [
            " In this paper, the 2-pole high-speed motor topologies with toroidal windings and different stator slot numbers, as well as their winding factors are described in section II. Then, the optimal designs of these motors are analyzed and validated by the finite element analysis (FEA) in section III. Section IV compares the electromagnetic performance of these motors with different slot numbers. Finally, section V is the conclusion. II. MOTOR TOPOLOGIES The topologies and specifications of four 2-pole highspeed (110krpm) PM motors with toroidal windings and different slot numbers are shown in Fig.1 and Table II. In order to reduce the electromagnetic and switching losses, the pole number is chosen as 2 so that the operating frequency is minimum. The rotor is surface-mounted with a 2-pole diametrically magnetized NdFeB magnet and the shaft employs magnetic material. The stator is also equipped with outer teeth which do not contribute to the magnetic circuit, but mainly provide the mechanical support. To ease winding, the stator employs straight teeth without tooth-tips. Further, due to the limitation of small size, the slot number is chosen as 3, 6, 9, and 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002481_00207543.2016.1259669-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002481_00207543.2016.1259669-Figure10-1.png",
        "caption": "Figure 10. Details of flow-thermo-structure BCs for \u2018take-off\u2019.",
        "texts": [
            " In this paper, objectives of optimisation are: (1) The minimum mass (W) should be less than 80 kg; (2) The minimum radial deformation (D) should be less than 1.95 mm. Constrain conditions of optimisation are: (1) The number of turbine blades is 68, each blade\u2019s mass is 0.144 kg and height is 56 mm; the distance from blade centroid to rotating axis is 298 mm; (2) The maximum allowable stress of the material is 798 MPa; (3) The coupled flow-thermo-structure BCs for working state \u2018take-off\u2019 are depicted in Figure 10(a). In Figure 10(a), \u2018\u2191\u2019 represents equivalent pressure; \u2018~\u2019 indicates heat transfer with mainstream gas; \u2018 \u2019 denotes zero displacement. Four curves along the border are cooling flows whose details are described by Figure 10(b). Parameters of these BCs are tabulated in Table 1. H1, H2, H4, H6, W2, W3, W4, W6, R3 and R4 are selected as factors for DOE. The span of each factor is summarised in Table 2. W1, W5, W7, H3, H5, R1, R2, R5 and R6 are set as 45.0, 42.0, 60.0, 195.0, 140.0, 280.0, 105.0, 5.0 and 6.0, respectively. With the help of historical knowledge, a mixed-level design is generated with (W2, W3, W4, H2, R3, R4) at three levels and (W6, H1, H4, H6) at two levels, Taguchi design L36(3^6*2^4) is constructed and corresponding responses are summarised in Table 3, maximum 1st principal stresses of these samples are less than 740 MPa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure9-1.png",
        "caption": "Figure 9. Geometry interference model of wheel/rail profiles.",
        "texts": [
            " Methods to calculate the wheel/rail static contact stress Because the wheel and rail surfaces are composed of different curved surfaces, the Hertz contact theory cannot be directly used to calculate the contact stress. Take one contact point M(xM, yM) for instance, its coordinates are the contact spot\u2019s center coordinates when actual deformation occurs, see Figure 8. It is assumed that there is a profile dividing point N near the point M, moreover, point N is in the contact spot when deformation occurs. On the basis of critical contact, the wheel profile moves downwards for a distance of y, and then the actual deformation at point M is replaced by geometry interference of the wheel/rail profiles, shown in Figure 9. By solving profile equations, the boundary points M1(x,y) and M2(x,y) of the interference area can be obtained, and the equation of a straight line between the two points can be written as y \u00bc ux\u00fe c. By referring to Houpert\u2019s idea of a slicing technique that is used to calculate roller/race contact19, the entire wheel surface is discretized into a number of slices along the x-direction. The width of each slice is x, as shown in Figure 10. The smaller the x is, the more accurate will be the calculation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000084_tie.2019.2920610-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000084_tie.2019.2920610-Figure2-1.png",
        "caption": "Fig. 2. Vector diagram of WRIM with (a) observer on stator and (b) observer on rotor (stator flux oriented).",
        "texts": [
            " From (5), the dq axes stator and rotor current ripples are approximately equal and opposite, since the term in the denominator (1 +\u03c3s) or (1 +\u03c3r) can be fairly approximated to 1. The simplified expressions for stator and rotor current ripples in arbitrary reference frame are as follows: i\u0302ds = \u2212i\u0302dr \u2248 1 Ll (\u03c8\u0302ds \u2212 \u03c8\u0302dr), i\u0302qs = \u2212i\u0302qr \u2248 1 Ll (\u03c8\u0302qs \u2212 \u03c8\u0302qr) (6) where, \u03c3Ls = \u03c3Lr \u2248 Lls + Llr = Ll. Therefore, to estimate the stator and the rotor current ripples, the estimated rotor flux ripples, in (D2-Q2) frame, are transformed to stator flux frame (D-Q) using the speed independent transformation in [9]. Fig. 2 shows the vector diagram of WRIM for subsynchronous and supersynchronous modes of operation of DIWRIM as seen from the stator and the rotor sides [9]. It can be observed that the phase sequence of the rotor voltage vector (Vr) changes as the machine operation moves from subsynchronous to supersynchronous mode [Fig. 2(b)]. Due to the change in the phase sequence, the Q2-axis flux ripple undergoes a change in the sign. Also, it is known that the magnitude of the supply frequencies in supersynchronous mode of operation are equal owing to (1). In supersynchronous mode, the transformation is independent of \u03c9et and thus, facilitates the open-loop estimation of current ripple [9]. Therefore, the transformed rotor flux ripples in (D1-Q1) frame can be written as: \u03c8\u0302D1r = \u03c8\u0302D2r, \u03c8\u0302Q1r = \u2212\u03c8\u0302Q2r (7) The transformation of rotor flux ripples from (D1-Q1) to (DQ) frame is simple using the load angle (\u03b4) as in [9]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002765_00207721.2020.1716277-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002765_00207721.2020.1716277-Figure2-1.png",
        "caption": "Figure 2. Schematic representation of EMLS.",
        "texts": [
            "8]T and select \u03b5 = 0.0165. Since there is no restriction on the timevarying rates of \u03b82(t), we set uncertain parameters from slow-varying rates to fast-varying rates such that (i) Case 1: \u03b4m(t) = 0.31 cos t and (ii) Case 2: \u03b4m(t) = 0.31 cos 1000t. Figure 1 shows that our controlmethod can treat fast-varying parameters as well. In this section, we will show the effectiveness of our control scheme via the vertical air gap control of an electromagnetic levitation system(EMLS). The singleaxis model of EMLS in Figure 2 is described by the following nonlinear dynamic equationwith the uncertain mass variation (Lee et al., 2000; Sinha, 1987): x\u03071 = x2 x\u03072 = \u2212\u03bc0N2A 4m(t) ( x3 x1 )2 + g x\u03073 = x2x3 x1 \u2212 2R \u03bc0N2A x1x3 + 2x1 \u03bc0N2A V (43) where x1 is the vertical air gap, x2 is the vertical velocity, x3 = i is the magnet current, V is the control input in voltage,m(t) = m + \u03b4m(t) is the time-varying total mass, m = 300 kg is the nominal mass, \u03b4m(t) is the uncertain mass variation, N = 660 is the number of turns of the coil wrapped around the magnet, A = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003662_s11668-020-01078-3-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003662_s11668-020-01078-3-Figure1-1.png",
        "caption": "Fig. 1 (a) Orientation of specimens from unidirectional fiber plate, (b) Schematic of modified Arcan test specimen and loading fixture",
        "texts": [
            " The curing cycle of Araldite 1092 epoxy was carried out at temperature 60 C, and time 20 min, the hand lay-up samples cured was performed through microwave. There was no elevated temperature post-curing of the sample performed. The mechanical properties of the E-glass fiber and the epoxy are presented in Table 1. Small size of the Arcan specimen allows specimens to be cut in all directions, so that any anisotropy in shear can be determined. The specimen width is the critical dimension, which in present work was 6 mm. A unidirectional 4- mm-thick plate of (GFRE) was laminated, and specimens were sliced from it in various directions, as shown in Fig. 1a. The axis connecting the notches is the axis for shear loading to determine the shear moduli G12 and G21. The select of G12 (fibers aligned with load direction) and G21 (fibers perpendicular to the load direction) specimens was motivated by the observations of the behavior of Iosipescu-type specimens [9]. The modified Arcan test specimen and loading fixture (composed of a pair of grips) are shown in Fig. 1(b). The specimen is tied to the fixture by three pins at each end. The specimen is loaded by pulling apart grips of the fixture at a pair of grip holes on the opposite sides of a radial line. By different loading angles a (a = 0 , 15 , 30 , 45 , 60 , 75 , and 90 ), all modes conditions starting from pure mode I to pure mode II can be created and tested [17]. Two groups of angle-ply ([0]4)s and ([90]4)s glass-fiberreinforced epoxy (GFRE) composite. The manufacturing route used for preparing the composite laminates is the hand lay-up technique",
            " The first template with zero oriented fibers is placed on the resin, then rolling displaces the air outward. When the ply is fully impregnated, the bundles are loosening from the template. The impregnation is repeated with alternate layers of resin and fibers, until the laminate is constructed. When the last ply is impregnated, it is covered by aluminum foil, and a load of about 5 kg is uniformly distributed on the glass plate. Then, a milling machine is used to cut the bonded plate into proper sizes. The specimens are cut with a diamond wheel and machined to dimensions, as shown in Fig. 1b. All tests were performed at room temperature using a crosshead displacement speed of 1 mm/minute. The specimens were loaded until specimen failure occurred. In laminate composite, the direction of fiber made angles 0 and 90 with the load direction and the fiber volume fraction (Vf) is 69%. The laminate order is shown in Fig. 2, while the elastic properties of unidirectional E-glass/epoxy lamina are shown in Table 2. An important point can be noted that the strain fields have been measured via Digital Image Correlation System composed by two video cameras to record pictures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001878_tmag.2015.2489701-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001878_tmag.2015.2489701-Figure2-1.png",
        "caption": "Fig. 2. Vertical view of two turns in the Cartesian coordinate system.",
        "texts": [
            " Normally, the orthogonal winding method (OGWM) or the orthocyclic winding method (OCWM) is used for wrapping a coil (Fig. 1). An analytical formula presented by Koch (K1-formula) in [2] can be used for OGWM with an air gap h between adjacent layers by calculating the electric energy stored between two layers. It is assumed that the wire is considered as an equipotential surface and electric flux lines in coating, and air gap regions are approximated by straight lines. A vertical view of two turns in adjacent layers is shown in Fig. 2. As the shape of the coil is arbitrary, only the layer-to-layer capacitance per unit length Cl_K 1 is then considered, and it can be expressed as follows [2]: Cl_K 1 = (2/3)n\u03b50 1 \u2212 \u03b4 \u03b5Dr ( F + 1 8\u03b5D ( 2\u03b4 r )2 1 1 \u2212 \u03b4 \u03b5Dr G ) with F = \u03b2\u221a \u03b22 \u2212 1 arctan (\u221a \u03b2 + 1 \u03b2 \u2212 1 ) \u2212 \u03c0 4 G = \u03b2(\u03b22 \u2212 2) (\u03b22 \u2212 1)(3/2) arctan (\u221a \u03b2 + 1 \u03b2 \u2212 1 ) \u2212 \u03b2 2(\u03b22 \u2212 1) \u2212 \u03c0 4 \u03b1 = 1 \u2212 2\u03b4 \u03b5Dr ; \u03b2 = 1 \u03b1 ( 1 + h 2r ) (1) 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www",
            " Accordingly, the layerto-layer capacitance per unit length Cl_M2 of OCWM is expressed as follows: Cl_M2 = n\u03b50 arctan ( ( \u221a 3\u22121)2(K+2) 2 \u221a K 2+2K ) 1.5 \u221a K 2 + 2K with K = 1 \u03b5D ln ( r r0 ) . (2) We can note that (2) is derived under the assumption that there is no air gap between adjacent windings, and the integration path of 60\u00b0 is satisfied for OCWM. Hence, a new formula for OGWM (M1-formula) can be deduced from (2) regarding the air gap h and an appropriate integration path of \u03b8 from \u221290\u00b0 to 90\u00b0 (Fig. 2). Then, Cl_M1 for OGWM can be stated by the following expression: ClM 1 = n\u03b50 arctan ((\u221a K + h 2r + 2 ) /\u221a K + h 2r ) 1.5 \u221a( K + h 2r )2 +2 ( K + h 2r ) . (3) As the formulas described above are derived under different physical insights, the accuracy of both formulas for OGWM should be investigated. Moreover, only the rectangular area denoted with dotted line shown in Fig. 2 is considered, because the fringing effect cannot be considered. It is clear that the approximation of electric flux lines by using straight lines is only validated if two turns are close to each other, and the angle \u03b8 is in the range of \u2212\u03c0/6 to \u03c0/6 [3]. That means, for the angle \u03b8 out of that range, the electric flux lines should be considered as arcs instead of straight lines. As a result, the approximation leads to an overestimation of the capacitance calculated by (3), because the equivalent distance in the air gap is smaller than its actual length, which results in a higher capacitance in the air gap Cair according to PPCF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure11-1.png",
        "caption": "Figure 11. Velocity field contours for the modified body geometry along the X-axis.",
        "texts": [],
        "surrounding_texts": [
            "An optimization calculation was carried out to assess the effect of the hull geometry. The original body geometry (see Figure 1) has been changed by moving the top down 70 mm - Figure 9. Figures 10 - 12 show the contours of the velocity fields and the velocity vectors for the modified housing shown in Figure 9. The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 31.2 m/s. For this geometry, a more even distribution of the air velocity fields is observed. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.2-1.png",
        "caption": "FIGURE 14.2 Model of a system subjected to shock input. (a) Cantilever structures subjected to shock input at support. (b) Equivalent single degree of freedom model.",
        "texts": [
            " In the present text, we will assume that students have some background in vibration analysis of systems. In particular, we will model systems as single-degree-of-freedom linear spring\u2013mass systems. This leads to a second-order linear differential equation (DE) whose closed-form solution is available (Chopra, 2007). It may be worthwhile for the students to briefly review the material on the solution of linear DEs in appropriate textbooks. To demonstrate the formulation and the solution process, we consider the cantilever structure shown in Fig. 14.2. The data for the problem and various notations used in the figure are defined in Table 14.11. The structure is a highly idealized model of many systems that are used in practice. The length of the structure is L and its cross-section is rectangular with its width as b and its depth as h. The system is at rest initially at time t = 0. It experiences a sudden load due to a shock wave or other similar phenomenon. The problem is to control vibrations of the system such that displacements are not too large and the system comes to rest in a controlled manner",
            " We will not discuss detailed design of the control force\u2013generating mechanisms, but we will discuss the problem of determining the optimum shape of the control force. II. NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION The governing equation that describes the motion of the system is a second-order partial DE. To simplify the analysis, we use separation of variables and express the deflection function y(x, t) as: \u03c8=y x t x q t( , ) ( ) ( ) (14.58) where \u03c8 (x) is a known function called the shape function, and q(t) is the displacement at the tip of the cantilever, as shown in Fig. 14.2. Several shape\u2013functions can be used. however, we will use the following one: \u03c8 \u03be \u03be \u03be( )= \u2212 =x x L ( ) 1 2 3 ;2 3 (14.59) Using kinetic and potential energies for the system, \u03c8 (x) of Eq. (14.59), and the data in Table 14.11, the mass and spring constants for an equivalent single-degree-of- freedom system shown in Fig. 14.2 are calculated as follows (Chopra, 2007): Mass: \u222b \u222b \u03c8= =     =my t dx m x dx q t mq tkinetic energy 1 2 ( ) 1 2 ( ) ( ) 1 2 ( ) L L 2 0 2 0 2 2 (14.60) where the mass m is identified as: \u222b \u03c8 ( )( )= = = =m m x dx mL( ) 33 140 33 140 1.56 1.0 0.3677 kg L 2 0 (14.61) y(x,t)=\u03c8(x)q(t) \u03c8(x)=123\u03be2 \u2212\u03be3; \u03be=xL kinetic energy =12\u222b0Lm\u00afy2(t)dx=12\u222b0Lm\u00af\u03c82(x)dx- q2(t)=12mq2(t) m=\u222b0Lm\u00af\u03c82(x)dx=33140m\u00afL=33 1401.561.0=0.3677 kg m\u00af II. NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION \u222b \u222b \u03c8= \u2032\u2032  = \u2032\u2032    =EI y x dx EI x dx q t kq tstrain energy 1 2 ( ) 1 2 ( ( )) ( ) 1 2 ( ) L L2 0 2 0 2 2 (14",
            " Depending on the starting point, the number of iterations to converge to the final solution varies between 20 and 56. Constraints on the terminal displacement and velocity of Eqs. (14.68) and (14.69) are active with the normalized lagrange multipliers as (6.395E \u2212 02) and (\u22121.771E \u2212 01), respectively. The control force is at its lower limit for the first 22 grid points and at its upper limit at the remaining points. It is interesting to compare the three formulations for the optimal control of motion of the system shown in Fig. 14.2. Table 14.12 is a summary of the optimum solutions with the three formulations. All of the solutions are obtained with 41 grid points and u(t) = 0 as the starting point. For the third formulation, T = 0.04 s is used as the starting point. The results in Table 14.12 show that the control effort is the greatest with the first formulation and the least with the second one. The second formulation turns out to be the most efficient as well as convenient to implement. by varying the total time T, this formulation can be used to generate results for formulation 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001040_j.mporth.2016.05.014-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001040_j.mporth.2016.05.014-Figure6-1.png",
        "caption": "Figure 6 In three dimensions, a rigid body such as the vertebra has six degrees of freedom. Relative to another object (such as the adjacent vertebra), it can rotate about and translate along any of the axes shown.",
        "texts": [
            " In general, in two dimensions, three coordinates are required to fully define the position of a rigid body: two define the x and y coordinates of a specific datum point on the body and the third defines the orientation of the body relative to that point. The body may translate, which will alter the x and y coordinates of the datum point, or rotate, which will alter the orientation coordinate. Each coordinate is known as a degree of freedom. In three dimensions, there are six possible degrees of freedom: three translational along three perpendicular axes and three rotational about each of the three axes (Fig. 6). Relative motion at the joint surfaces In the diarthrodial joints, the relative motions of the bones are constrained by the geometry of the joint surfaces and action of the ligaments and muscles spanning the joint. When the two 2005 Elsevier Ltd. All rights reserved. joint surfaces remain in contact, they may move relative to each other by rolling or sliding. Figure 7 shows the simple case of a circular wheel on a flat surface. Rolling occurs when there is no relative velocity, that is, no slip, between the two contacting points and the ICR is located at the point of contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002401_aim.2016.7576930-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002401_aim.2016.7576930-Figure4-1.png",
        "caption": "Fig. 4. Slave forceps\u2019 manipulator",
        "texts": [
            " Let the two ends of the needle be the tail GP1 connected to suture and the tip of needle GP2\uff0eStarting with the needle held in the Slave A1\u2019s gripper at GP1\uff0e Master-slave system consists of one master and two slave forceps\u2019 manipulator (Slave A1 and A2)\uff0eFigure 3 shows the flow chart of an automation of suture task with Slave A1 and A2 which are controlled by a Master\uff0eAt initial state\uff0cMaster is connected with Slave A1 by bilateral control and insert needle to tissue with Slave A1 manually\uff0eAfter this step\uff0c Slave A2 recognizes the insertion automatically and Slave A2 starts approaching to GP2 automatically\uff0eThen\uff0cSlave A1 hands-off the needle to Slave A2 automatically\uff0eIn this state\uff0c Master\uff0cSlave A1 and A2 are connected with multilateral control to send back both estimated external force of Slave A1 and A2\uff0eAt the end\uff0cMaster connects with Slave A2 by bilateral control and pull the needle from tissue with Slave A2 manually\uff0eTo achieve this proposed procedure\uff0cwe need to install followings to control system\uff0e \u2022 Automation of needle inserting recognition \u2022 Automation of Slave A2 approach to GP2 1) Master device: We used Geomatic Touch R\u20dd(3D Systems\uff0cInc.)[4] for haptic device to control slave forceps\u2019 manipulator globally\uff0eThis device has 6-DOF in position and posture\uff0cand translational 3-DOF to present force to operator\uff0e 2) Slave device: We used slave forceps\u2019 manipulator as shown in Figure 4[5]\uff0eThis system uses the pneumatically driven forceps manipulator with flexible joint\uff0eThis manipulator can operate 7 DOF in total such as up/down\uff0cright/left\uff0c zoom in/out\uff0croll\uff0cand grip\uff0eBesides\uff0cusing high backdrivability of pneumatic actuators\uff0cit is possible to estimate the external force (about 0.3-5[N]) on the forceps tip\uff0e We established a threshold for Slave A2 to start autonomous approach to grasping point of the needle GP2 and selected the value of threshold Tcn experimentally \uff0e Figure 5(a) shows the experimental environment\uff0eWe conducted this experiment with a curved needle which is used in surgical task and a phantom resembling the human bile duct \uff08Wetlab\uff0cbile duct model training kit: \u03a68[mm]\u00d760[mm]\uff09\uff0e In this experiment\uff0cwe measured two factors to utilize for a threshold\uff1amagnitude of force and rotation angle of Slave A1\uff0eWe defined the multiplication of these two factors as following\uff0e1 Wcn = \u221a f 2 ext"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001575_1464419314540901-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001575_1464419314540901-Figure3-1.png",
        "caption": "Figure 3. The ideal roller bearing model. x is the revolving speed of rotor; VA and VB are speed of the contact points between the roller and the races; Vc is revolution speed of roller; rb and Rb are radius of inner and outer race respectively; j is the revolving angular of jth roller.",
        "texts": [
            " With the help of Hamilton principle and finite element, the systematic dynamic equation can be described as M \u20acq\u00feG _q\u00fe K q \u00bc Fb \u00fe Fu \u00fe Fg \u00f01\u00de where M and G respectively stand for the mass matrix and gyroscopic matrix of the rotor after considering the disk, K is the stiffness matrix of the rotor, Fb is the vector of nonlinear supporting forces of bearings, Fu is the periodic exciting force (i.e. unbalanced excitation with the same phase as the rotating speed) vector acting on the rotor, and Fg is the gravity vector of the rotor, and q\u00bc (x, y, , \u2019) is the displacement vector of the rotor. The roller bearing model considered here has equispaced rollers rolling on the surfaces of the inner and outer races as shown in Figure 3. Elastic deformation Figure 1. Roller bearing\u2013rotor model. at WEST VIRGINA UNIV on June 21, 2015pik.sagepub.comDownloaded from would be supposed to arise at the contact lines between the roller and races after loading. While the bearings are lubricated with adequate preloading or interference fitting, one can assume that the prevailing regime of lubrication is elastohydrodynamic, which means that the Hertzian line contact theory20 could be an approximation for the restoring force of elastic deformation between roller and raceways",
            "22 Through the study of Harris and Kotzalas,23 the kj could be calculated by kj \u00bc 1 \u00f01=ki\u00de 9=10 \u00fe \u00f01=ko\u00de 9=10 10=9 \u00f03\u00de where ki is the stiffness between the jth roller and inner race, similar to ko but corresponding to the outer race. Then, the common Palmgren24 equations could help to establish the value of ki and ko, which was also adopted by Gupta.25 The rotation angle of the jth roller at moment t is j \u00bc 2 Nb \u00f0 j 1\u00de \u00fe !ct \u00f04\u00de where Nb is the total amount of rollers and xc is the angular velocity of the cage. According to the model in Figure 3, the angular revolution velocity of the cage, which equals the revolution velocity of roller, can be expressed as !c \u00bc !rb rb \u00fe Rb \u00f05\u00de Then the VC frequency is given by !vc \u00bc !c Nb \u00bc !rb rb \u00fe Rb Nb \u00bc ! BN \u00f06\u00de After taking into account the radial clearance , the elastic deformation of the jth roller can be calculated as j \u00bc x cos j \u00fe y sin j \u00f07\u00de where x and y are the displacements of inner race center respectively in horizontal and vertical directions. From equations (2) and (7), the contact force between the jth roller and the races is obtained as Q j \u00bc kj\u00f0 j\u00de 10=9, j 4 0 0, j40 ( j \u00bc 1, 2, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000730_20130904-3-fr-2041.00089-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000730_20130904-3-fr-2041.00089-Figure4-1.png",
        "caption": "Fig. 4. View of the AC-ROV with the reference frames (xiyizi: earth-fixed frame, xbybzb: body-fixed frame).",
        "texts": [
            "9 if |e| \u2264 0.1r The obtained simulation results are shown in Fig. 3. The system output is displayed for both versions of the controller. The original L1 adaptive controller (dotted blue line) exhibits a clear time lag that becomes insignificant when the proposed proportional extension is added (solid black line). This example illustrates the benefits of extending the original architecture proposed in Hovakimyan and Cao [2010] as explained in this section. UNDERWATER VEHICLE The AC-ROV submarine (cf. Fig. 4) is an underactuated underwater vehicle. The propulsion system consists of six thrusters driven by DC motors controlling five degrees of freedom. Four horizontal thrusters control simultaneously translations along x and y axes and rotation around the z axis (yaw angle). The two horizontal thrusters denoted \u2019Thruster 1\u2019 and \u2019Thruster 2\u2019 on Fig. 3 control depth position and pitch angle. The roll angle is unactuated but remains naturally stable due to the relative position of buoyancy and gravity centers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002827_j.optlaseng.2020.106065-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002827_j.optlaseng.2020.106065-Figure15-1.png",
        "caption": "Fig. 15. The state of the particle and the surrounding fluids at 29 \ud835\udf07s: (a) where the powder particle is located; (b) speed distribution of the particle and the surrounding fluids; (c) the speed distribution of the pool around the particle.",
        "texts": [
            " 14 at moving time of 18 \ud835\udf07s and has an immersion depth about one-half of ts own diameter in the melt pool, the powder particle speed decreases o ~1.86 m s \u2212 1 , which is ~74.4% of its initial speed. As can be seen rom Fig. 14 (c), the fluctuation zone of the melt pool increases in depth urther, and the fluctuation area at the gas-liquid interface increases ignificantly compared to that at 12 \ud835\udf07s. The peak speed of the melt ool appears around the particle, which is ~1.86 m s \u2212 1 , and the speed radient was reduced compared to those at 6 \ud835\udf07s and 12 \ud835\udf07s. Fig. 15 shows the state of the powder particle and the melt pool at 9 \ud835\udf07s as the particle is immersed in the melt pool to a depth of twohirds of its diameter. It can be seen that the liquid level of the melt ool around the particle is lower than that of the mean level of the gasiquid interface ( Fig. 15 (a)), and the speed of the particle is ~1.71 m s \u2212 1 , hich is ~68.4% of its initial speed. The area of the fluctuation zone f the melt pool increases significantly at the gas-liquid interface, and he peak speed is generated around the particle, which is ~1.71 m s \u2212 1 , hile the expansion speed of the outermost periphery of the fluctuation one decreases to less than 0.3 m s \u2212 1 . Fig. 16 shows the powder particle and melt pool state at 53 \ud835\udf07s. The owder particle has completely immersed in the melt pool, but the melt ool is not completely closed. A large air gap is generated between the pper melt pool and powder particle, and the length of the gap is sigificantly expanded relative to Fig. 15 . The gas-liquid interface of the ool is significantly disturbed, as the reverse height difference is obious. It can be seen from Fig. 16 (b) that after the powder particle is ompletely immersed in the gas-liquid interface, the speed of the partile is decreased to ~1.27 m s \u2212 1 , which is ~50.8% of its initial speed. he fluctuation zone at the gas-liquid interface increases significantly, nvolving the entire calculation zone in this work. The peek speed of he melt pool fluctuation zone is ~1.27 m s \u2212 1 and the speed gradient s decreased significantly compared to that in the initial stage (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001012_issnip.2014.6827628-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001012_issnip.2014.6827628-Figure2-1.png",
        "caption": "Fig. 2. Bending causes strain.",
        "texts": [
            " Several gratings can be multiplexed in one fiber by assigning different \u201dbase wavelengths\u201d to each grating, reading the signal of each grating at the same time (wavelength division multiplexing, WDM). The principle is also shown in Fig. 1. One disadvantage for our purpose is the temperature dependence of the FBGs, as temperature also influences the reflected wavelengths. This temperature dependence can be compensated by adding one fiber which only measures temperatures, but so far, this aspect has been left out for our experiments as we only work under laboratory conditions. The idea of shape sensing is to measure the strains caused by bending, see Fig. 2. With at least three strain measurements in one cross-section, the curvature and local bend direction can be calculated. This means that three fibers with several gratings have to be guided along the instrument in parallel. The shape can then be reconstructed using different theories, depending 978-1-4799-2843-9/14/$31.00 \u00a9 2014 IEEE on the application: linear or cantilever beam theory for biopsy needles [2], [3] or using Frenet frames for endoscopes or other flexible instruments [4], [5]. While the mathematical theory for those types of shape sensors seems to be well analyzed, the problem of attaching the fibers to the instrument remains"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003582_biorob49111.2020.9224440-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003582_biorob49111.2020.9224440-Figure1-1.png",
        "caption": "Fig. 1. (a) The novel foot interface. (b) Isotonic foot interface [23]",
        "texts": [
            " \u2022 Intuitiveness: The mapping between the foot and the robotic arm movement should be natural and direct, which accords with the user\u2019s habits and is easy to learn [31]. i.e. sliding the foot to the right actuates movement to the right for the robotic arm. The mapping should also reduce the cognitive load for surgeons. \u2022 Ergonomics: All foot commands should be within the comfortable range of motion for the surgeon. The physical burden should be minimised during use. Also, the user should be able to step on the FI stably. The novel FI (480X360X145 mm3) consists of a hollow, sunken box and a central foot pedal (Figure 1(a)). Unlike Huang\u2019s design [24], our interface deploys the single DoF control strategy, which is a common strategy for commercial foot interfaces [9], [22], and two-grade speed selection. We hypothesize that manipulating coupled DoF simultaneously in surgery applications makes predicting the final position of the end effector more difficult and puts patient safety at risk. Moreover, keeping a stable speed in the constantly changing rate control is challenging and may create extra physical and mental burden to the user",
            " A flat aluminum spring beneath the pedal is used to provide the reverse torque for the pedal rotation, the stiffness of which can be customized easily by adjusting the thickness of the flat spring or the number of sheet layers. To satisfy the requirements for safety, all DoF are mechanically decoupled. Mechanical decoupling means each of the movements on the four inclined surfaces and the two modalities of pedal rotation are mapped to a distinct DoF. Therefore the user always actuates one DoF at a time and cannot actuate any other DoF. This mechanical structure 1In pilot tests, we tried angle setup from 10 degrees to 25 degrees with a step of 5 degrees addresses the coupling issue in the isotonic interface (Figure 1(b)), which may cause the user to send a wrong command since the movement to rotate the pedal (control Roll) can force the pedal to slide (control Pitch) in the rail as well. On the software side, the proposed FI uses the single DoF control strategy i.e. only one DoF can be actuated at a time, and all the other DoF are locked through the controller. Besides, the actuation mechanism has a 0.4 second buffering. If the user moves their foot back to the central pedal within 0.4 seconds, the last command will not be actuated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000529_00423114.2015.1046462-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000529_00423114.2015.1046462-Figure2-1.png",
        "caption": "Figure 2. Mechanical layout of a roller rig for tests on a single wheel set.",
        "texts": [
            " The results of the proposed methodology are assessed by means of numerical experiments performed using a mathematical model of the roller rig, while the use of this methodology in real tests on a full-scale roller rig is envisaged as a future step of the research. This paper considers the use of the BU300 roller rig installed at Lucchini R.S. (shown in Figure 1). Other implementations of the same concept are the RaSSP roller rig owned by DB Systemtechnik, a roller rig by KRRI [13] and the one at CARS.[5] More information about the physical systems and simulation technologies can be found in references [6,14,15]. The mechanical layout for a roller rig of this type is shown in Figure 2: it consists of a roller with two wheels driven by a motor. A wheel set is mounted on the top of the roller and connected through primary suspensions to a transversal beam representing one half of the bogie frame. Figure 1. The BU300 roller rig at Lucchini R.S. D ow nl oa de d by [ U ni ve rs ity o f L et hb ri dg e] a t 1 5: 23 2 8 Se pt em be r 20 15 1334 B. Liu and S. Bruni The roller rig actuation system consists of multiple modules accomplishing different tasks: (1) a vertical actuation module, consisting of two actuators used to apply a desired combination of vertical forces on the two sides of the half-bogie; (2) the lateral actuation module, consisting of a single actuator which is used to apply a lateral force on the half-bogie; (3) the yaw actuation module, which is used to control the yaw rotation of the half-bogie. This module may consist of a variable number of actuators (4 in the scheme of Figure 2), depending on the specific mechanical layout adopted in the design of the rig; (4) a traction module, driving the rollers according to a desired speed profile. Modules 1 and 2 are typically operated in the force control mode, whereas module 3 is normally operated in position control and module 4 in speed control. Based on the authors\u2019 knowledge, at least for the BU300, KRRI and CARS benches, the references of actuation modules 1, 2 and 3 can be defined in various ways which include manual control by a human operator, the execution of a sequence of simple waveforms such as ramps, holds, sinusoids or generic time histories that are generated externally and fed to the control system prior to the execution of the test. In this paper operation based on externally generated reference time histories is assumed. In Figure 2, which corresponds to the existing implementations of roller rigs for a single wheel set, the two rollers are rigidly connected and actuated to rotate at the same angular speed. In this paper, besides this base arrangement of the traction module, a modified design is considered which includes a differential gear with the purpose of having slightly different angular speeds of the two rollers. It should be noted that such an arrangement is at present not used in any of the existing roller rigs mentioned above, however it is used in some scaled and full-scale roller rigs for tests on bogies or vehicles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003449_j.measurement.2020.108458-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003449_j.measurement.2020.108458-Figure2-1.png",
        "caption": "Fig. 2. Hardware setup. 2(a) Instrumented shoe prototype with a foot-to-ground angle module mounted laterally, 2(b) Top and side view of the foot-to-ground angle module that consists of a pair of IR sensor S1 and S2.",
        "texts": [
            " Therefore, we utilized only sagittal plane FGA because of its large variation in amplitude dynamics, which accounts for a large scope for terrain classification. The assumption of not considering the coronal plane angle benefited in two ways: \u2212 1) it avoids the use of the additional sensor for coronal plane FGA measurement, 2) since sagittal and coronal plane FGA are interdependent [32] therefore, it reduces the calculation complexity for obtaining sagittal plane FGA. The top and the side view of the FGA module is shown in Fig. 2(b). It consists of two IR sensors placed at a fixed distance from each other with a partition in between to avoid any interference between the IR beam of the two sensors. IR sensors were calibrated by recording voltages (in volts) corresponding to every 0.5 cm increment in the distance of the sensor from the flat surface. Then, a fivedegree polynomial equation was obtained using a curve fitting toolbox in MATLAB 2017a (MathWorks Inc., USA, academic license). The IR sensors used has a minimum threshold below which the measurement is not accurate (as per the datasheet from the manufacturer [33]). Therefore, the sensors were installed at a certain vertical distance from the flat ground on the shoe. The IR sensors of range 10\u201380 cm (model no. 2Y0A21YK0F) were used, which was sufficient for the current application of FGA measurement. The instrumented shoe prototype with the FGA module installed laterally is as shown in Fig. 2(a). The data acquisition system consists of the Arduino micro board (ATmega32U4 8-bit microcontroller) and HC-05 Bluetooth module for one-way wireless serial communication between remotely located PC and the data acquisition system. The data acquisition circuit was powered using 9 V, 600 mAh rechargeable battery (from EBL, USA). The data was transmitted at a sampling rate of 86 Hz after it was exposed to the Kalman filter. A Kalman filter was employed for a better estimate of the measurement. The Kalman filter uses a probability-based approach to estimate the current value of the sensor using the current measurement and previous estimate",
            " The slope of the ramp considered in the present study is 5\u25e6, which agrees with the standard slope allowed in the buildings in different countries [53]. Although the FGA dynamics on the different sloped surface has not been investigated here, the authors believe that increasing the slope of the surface may further increase the classification accuracy due to significant change in ankle kinematics at higher slopes [54]. The sensor assembly offers the movement of angle module in the two directions as shown earlier in the Fig. 2. The movement of assembly in the vertical direction facilitates the use of the FGA module for varying shoe heights, whereas the movement in the horizontal direction facilitates the use of the FGA module for different foot lengths. Although the proposed method is novel and has certain merits compared to the prior methods, there are still a few limitations associated with the method. The major limitations are: 1) As the current method utilizes IR sensing-based FGA estimation, the IR sensor measurement is affected by the scattering of the IR beam because of the surface roughness and the contamination of the sensor lens with dust particles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003511_j.triboint.2020.106696-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003511_j.triboint.2020.106696-Figure6-1.png",
        "caption": "Fig. 6. Test rig of wear of brush seal.",
        "texts": [
            " The sliding distance was determined when the depth of the scar on the disc sample was similar to the variation of the inner diameter of the brush seal. The ball sample has a coating structure and the basis material of the ball was structural steel. The element types of the surface of the wear scar on the ball sample was shown in Fig. 5 and there was no Fe element. It was proved that the coating on the ball still remained after the test. The experiments are studied on the purpose of determination of the correction factor \u03c8 and validation of the mathematic model. All the experimental works were performed in a sealing test rig as shown in Fig. 6 which can be used to research all kinds of seals such as labyrinth seal, brush seal, finger seal and so on. In this paper, a brush seal was installed in position 1, and a labyrinth seal was installed in position 2 to form the HP-chamber as shown in Fig. 6 (a). The rotor was supported by two deep groove ball bearings and a roller bearing. A motor with the maximum speed of 10,000 rpm drove the rotor with the aid of a belt. The high-pressure air was just compressed air from the atmosphere. High pressure sealing air entered the high-pressure chamber from the middle and low pressure air leaked from both sides. Five air inlets guaranteed the circumferential uniformity of high pressure air. Three air inlets can be seen in Fig. 6 (b), and the other two were on the covered side. J. Fan et al. Tribology International 154 (2021) 106696 The instruments of the experiments include five parts, mass flow rate measurement, pressure measurement, temperature measurement, rotating speed measurement and wear of brush seal and rotor measurement. The first four parts were measured on the brush seal test rig and the measurement taps were around the seals. Magnify the area around the seals and the schematic diagram of measurement taps was shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000934_1.4024781-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000934_1.4024781-Figure7-1.png",
        "caption": "Fig. 7 Laser microscope with wide field of view and rotary table with the gasket. (a) Laser microscope with a gasket. (b) Layout of the laser beam and pixel data.",
        "texts": [
            " The laser microscope with wide field of view we developed is not suitable for examination of the surface profiles because there is a trade-off between the field of view and the resolution in the depth direction. As we attempted to expand the field of view of the laser microscope, the resolution in the depth direction was about 6 lm. The maximum roughness of the flange was 6.3 lm and that for the gasket surface was 0.3 lm. The flange and gasket surface roughnesses are shown in Fig. 6. 3.2 Observation Procedure Using a Laser Microscope With a Wide Field of View. Figure 7(a) shows a confocal laser scanning microscope with a wide field of view of 10 mm. A laser beam is scanned by a rotating flat mirror at a rotational speed of 9000 rpm and is focused on an area about 3 lm in diameter on the focal plane. A stepping motor-driven rotary table can be seen on Journal of Tribology OCTOBER 2013, Vol. 135 / 041103-3 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 03/10/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the left side of Fig. 7(a), on the center of which the copper gasket was set. In this configuration, all the contact marks formed on the gasket surface can be observed within a very short time. The resolution of the laser microscope was evaluated using a USAF resolution target consisting of bars organized into groups and elements (Edmund Optics). The resolution in the laser scanning direction was about 2 lm. Thus, the even narrower leakage paths on the gasket surface can be observed with this laser microscope. The laser microscope is constructed based on the principle of confocal microscopy as shown in Fig. 7(b). As a scanning lens, the fh lens is assembled based on a shrink fitter technique [14]; the sizes of the laser spots on the focal plane can be kept constant over a wide scanning width. This is why this laser microscope has a wide field of view. First, a laser beam emitted by a laser diode is collimated by a collimator lens. The wavelength of the laser is 650 nm. In addition, a linearly polarized collimated laser beam is transformed into a circularly polarized beam by a quarter wave plate. The laser beam is scanned by a rotating flat mirror at a rotational speed of 9000 rpm and passes through the fh lens unit to focus the beam on the focal plane at a laser spot with a diameter of about 3 lm over a scanning width of 10 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000033_aim.2019.8868339-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000033_aim.2019.8868339-Figure1-1.png",
        "caption": "Figure 1.Cut-off line.",
        "texts": [
            " Besides, the number of fatalities per vehicle-miles driven is higher at night than during daytime [2] and this is strongly related to the fact that during nighttime (or in a more general scenario under low visibility conditions), the risk of collision increases as the ability to detect dangers decreases. The most important elements for good detection at nighttime are the headlamps, easily described by the cut-off line, which divides the panorama in two parts: the bright area below that provides visibility to the driver and a dark area above that prevents glaring on oncoming traffic (Figure 1). When driving, the constant change in road conditions, together with the variable loads on the vehicle, have a direct effect on the inclination of the cut-off line and thus in the detection capabilities and safety : if the cut-off line rises above 1 Authors are with Roberval laboratory CNRS, FRE 2012, Sorbonne Universit\u00e9s, Universit\u00e9 de Technologie de Compi\u00e8gne CS 60319-60203, Compi\u00e8gne cedex, France. (Email: svenegas@utc.fr, frederic.lamarque@utc.fr, christine.prelle@utc.fr,). a certain height, then glare may be generated on other drivers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003256_1.c035739-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003256_1.c035739-Figure11-1.png",
        "caption": "Fig. 11 Multi-rigid-body dynamics model of landing gear.",
        "texts": [
            " For flexible body dynamics model, the derivation of dynamics equation is far more difficult than the multi-rigid-body dynamics equation due to the larger number of voxels. Hence, in this paper, the Siemens Virtual. Lab Motion multibody dynamics software is used to handle the multiflexible body problem, and to build the dynamics model of the 3-D retraction mechanism with hinge clearance and axis deviation. The linkage part models of landing gear are simplified while ensuring the inertial mass of mechanism. As shown in Fig. 11, the damper strut, wheel axle, torsion arm, and wheels are deemed as an integral rotating part andnamedmain strut,whereas all the remainingbars of the sidestay are separate dynamic components. Kinematic pair connection is established according to the linking relations between the structural parts shown in Fig. 12. The main strut is connected to the fuselage via two spherical hinges. Thenodes in the sidestay and the interconnections between the sidestay links are all rotation pairs, whereas the connection pairs betweenvariousmembers are all connected in the form consistent with the landing gear in an actual airplane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure23-1.png",
        "caption": "Fig. 23. Layering of antiparallelogram unit.",
        "texts": [
            " For instance, plane-symmetric Mode I can be transited to plane-motion Mode V through plane-symmetric Mode II or plane-motion Mode III or plane-motion Mode IV. In the design of prototype of the multiple-mode mechanism composed of four antiparallelogram units and four R joints, several factors should be taken into consideration. To make the mechanism achieve Singular configuration II and Singular configuration III, the links of antiparallelogram unit should be arranged in three layers as shown in Fig. 23 . To make two antiparallelogram units coincide, the R joints at the end of the antiparallelogram units need to be set offset angle g , and this offset angle g should be satisfied g = arccos(2 t / u ). However, due to the offset angle, the mechanism can only achieve the whole inward configurations and a part of the outward configurations. A prototype of the multiple-mode mechanism is fabricated with parameters l = 66.00 mm, l 1 = 93.34 mm. Several snapshots of the prototype in different modes are shown in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001179_s12541-014-0399-5-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001179_s12541-014-0399-5-Figure3-1.png",
        "caption": "Fig. 3 Design of fuel cladding and end cap",
        "texts": [
            " The control PC controls the welding head, shield gas, and three-axis LM guide according to the designated welding path and laser power. The welding position, path, and beads can be checked through the separated monitor whose image is transferred from the CCD camera. The developed fiber welding system is applied to test the welding performance of the nuclear fuel clad and the end cap made of Zircaloy4 (Table 1). A nuclear fuel rod is fabricated by putting nuclear fuel pellets, alumina pellets, and plenum into the fuel clad, and sealing the fuel clad with two end caps (Fig. 3). Their welding will be carried out in a circumferential direction by rotating the fuel clad with 0.087 rad/ sec of angular velocity. Preliminary experiments were carried out using CW mode and QCW mode at the focal length (180 mm) between focal lens and the workpiece. As shown in Fig. 4, sparks frequently fly up and the bead is not generated smoothly because some of the molten metal is burnt out or flows out of the bead. Because the peak power of a laser beam is configured at the focal length, the molten metal generated by the previous beam spot absorbs the beam energy at the overlapped area, and some of the molten metal is burnt out"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003568_10.0001659-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003568_10.0001659-Figure9-1.png",
        "caption": "Fig. 9. Control volume used to illustrate the conservation of momentum for a ball subject to the Magnus force. The large distances to the lateral control surfaces make pressure differences negligible (Ref. 25).",
        "texts": [
            " This phenomenon is referred to as the inverse Magnus effect because such a ball curves \u201cthe wrong way\u201d compared to what we are used to from properly designed sports balls. The asymmetric flow separation explaining the inverse Magnus effect is illustrated in Fig. 8. Beautiful flow visualization photographs capturing the Magnus effect and the inverse Magnus effect are presented in Ref. 23. The Magnus force as well as the inverse Magnus force can be understood by considering the conservation of momentum using the control volume illustrated in Fig. 9. The most intuitive control volume is thin in the flight direction and very long in the direction of the Magnus force.25 In such a control volume, the integrated pressures on the narrow upper and lower sides are approximately equal, such that the integrated sideways momentum flux of the flow entering and exiting the control volume in the front and the rear of the ball is equal and opposite to the Magnus force. Note that the flow entering the control volume has sideways momentum in the same direction as the Magnus force, and the flow exiting the control volume has sideways momentum in the direction opposite to the Magnus force as illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000617_9781118717806.ch5-Figure5.62-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000617_9781118717806.ch5-Figure5.62-1.png",
        "caption": "Figure 5.62 Solid and litz wires. (a) Solid wire. (b) Litz wire. Shaded areas indicate the skin depth",
        "texts": [
            " If a multistrand conductor is used, the overall cross-sectional area is spread among many conductors with a small diameter. For this reason, a stranded and twisted wire in the form of a rope results in a more uniform current density distribution than a solid wire. Moreover, litz wires are assembled so that each single strand, in the longitudinal development of wire, occupies all positions in the wire\u2019s cross section. Therefore, not only the skin effect, but also the proximity effect, is drastically reduced at high frequencies. Figure 5.62 shows a solid wire and litz wire; the shaded area indicates the skin depth. Litz wire has a lower AC resistance than a single round solid wire of the same cross-sectional area. It is important to position each individual strand in a uniform pattern moving from the center to the outside and back in a given length so that each strand takes up all possible positions in the cross section in order to reduce the proximity effect. Litz wire intended for higher frequencies requires more strands of less diameter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure6-1.png",
        "caption": "Fig. 6. Computational kinematic model of a PM with three active prismatic joints.",
        "texts": [
            " The Chebychev-Gr\u00fcbler-Kutzbach (CGK) formula was used to determine the mobility (m = 3) of the hydraulic TPM [23]. Experimental studies of the 3-RRPRR TPM with three IEHSDs revealed unexpected additional degrees of freedom. These come about when the hydraulic actuator piston rods are not blocked from rotating. In such situations, the kinematic mechanism of the TPM, instead of being rigid, behaves as if it had additional degrees of freedom. Therefore, a computational kinematic calculation model of parallel manipulator with three prismatic joints (hydraulic actuators) was introduced (Fig. 6). The kinematic model of parallel manipulator is useful for analyzing the EE trajectory in cases of slack and non-rigidity in rotary joints. The kinematic mechanism, after fulfilling certain conditions, will allow the obtaining of pure translation parallel manipulator movements. The computational kinematic model shown in the Fig. 6 corresponds to the 3-UPU architecture parallel manipulator [24, 25]. Since each U joint consists of two intersecting revolute R joints, each leg is equivalent to an RRPRR kinematical chain. The kinematic mechanisms of the 3-UPU and 3-RRPRR architecture parallel manipulators are already well known. For the purpose of kinematic analysis, two coordinate frames of the EE were defined: Position by coordinate: xp, yp, zp and orientation by angles: a, b, g. These relate to the RPY (Roll-Pitch-Yaw) rotation angles about the X, Y, Z axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure2-1.png",
        "caption": "Figure 2. Finite-volume partitioning.",
        "texts": [
            "225 ,  Viscosity \u2013 1.79 kg/(m\u2219s),  Acceleration of gravity 9.81 . Numerical-analytical algorithm was implemented using two transport equations for turbulent characteristics (k-\u03b5 model) [18]. To obtain the fields of air velocities in the housing of the final threshing device, this problem was solved numerically by the method of finite volumes [17]. To implement the numerical-analytical algorithm, the following conditions were set, boundary conditions: 1. Fan speed in rpm is set. 2. It section 1 (Figure 2), air is drawn in, the pressure is 0 Pa. Here and below, the pressure is specified as the difference between atmospheric pressure. 3. In section 3, Figure 2, the air flow is released, the pressure is 0 Pa. 4. In section 2 (Figure 2), the air flow rate is set to 20 m/s. 5. Sticking conditions are set on the other walls of the casing, the casing and the walls of the fan. All geometrical dimensions, drum rotation speed and air flow drawing speed were parameterized to make it possible to study the dependence of the sought parameters on these characteristics. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 Figures 2, 3 show a finite-volume partition of the internal air domain of a field machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure9-1.png",
        "caption": "Figure 9. Four types of closed-loop RRRS chains forming an isogram: (a) Type I (b, ; a)j(a, ; b), (b) Type II (b, ; a)j(a, ; b), (c) Type III (b, ; a)j(a, ; b), and (d) Type IV (b, ; a)j(a, ; b).",
        "texts": [
            " The six loops are movable with one DoF and they are symbolized by \u00f0b, ; a\u00dej\u00f0b, ; a\u00de, \u00f0a, ; b\u00dej\u00f0a, ; b\u00de, \u00f0b, ; a\u00de j\u00f0a, ; b\u00de; \u00f0b, ; a\u00dej\u00f0a, ; b\u00de, \u00f0b, ; a\u00de j\u00f0a, ; b\u00de, \u00f0b, ; a\u00dej\u00f0a, ; b\u00de The first two combinations have a minor interest. They characterize simply plane-symmetric RRRSloops in which the plane of symmetry is the plane containing the R axis, (O, k), and the S center, M. As a matter of fact, a plane symmetry does not alter the lengths and changes the angle signs. Moreover, the discrimination between (b, ; a)j(b, ; a) and (a, ; b)j(a, ; b) is not significant because the exchanges b$ a and $ can result from a change of notation. Henceforth, one can focus on the last four combinations shown in Figure 9, which can be described with the help of isograms. Bennett8 coined the word isogram to designate a non-planar quadrilateral having a line symmetry when he found out a new 4R linkage. An isogram can be viewed as a warped parallelogram. Any isogram is globally invariant in a half-turn around a straight line. Each side and each vertex has its homologous in the line-symmetry, which is said to be its opposite. The lengths of two opposite sides are equal. From a couple of opposite vertices, a segment of straight line can be drawn"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002769_ls.1499-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002769_ls.1499-Figure3-1.png",
        "caption": "FIGURE 3 Improved rolling bearing assembly including through driving shaft cooling\u2013heating",
        "texts": [
            " Later, a modification to the housing of the assembly allowed performing tests at controlled temperature as long as the stabilisation temperature was lower than the inherent operating temperature of the rolling bearing. The main challenge in the recent development of the modified Four-Ball machine was in the necessity of heating and cooling the assembly minimising at the same time the risk of interfering with the measurement of the frictional torque loss. In order to achieve that goal and retain the active temperature control capabilities, an alternative solution was found: The rolling bearing assembly was heated/cooled through the centre of the driving shaft of the Four-Ball machine. Figure 3 shows the proposed solution. The thermal fluid, flows down through the \u201c8-Coolant inlet\u201d that is floating in the centre of the driving shaft expanding into a \u201c6-Coolant chamber\u201d at its end. The fluid is then forced up the driving shaft through the clearance between the centre hole of the driving shaft and \u201c\u20186-Coolant inlet\u2019, the \u201c9-Coolant outlet\u201d. In this particular arrangement, the heat is extracted from the assembly through the driving element of the inner race of the rolling bearing (Key: 4, Figure 3). Heat is then exchanged through the interface between elements 4 and 10, Figure 3, to the thermal fluid that circulates in the \u201cCoolant Chamber\u201d. This way of transferring heat through the driving element of the inner race of the rolling bearing circumvents the need of having hoses attached to the external housing of the assembly that could potentially interfere with the friction torque measurement. Key Tsump Ttop 1- Rolling bearing outer race support (fixed); 2,9- Upper/lower locking pin; 3- Housing; 4- Rolling bearing; 5- O-ring; 6- Drive Spindle; 7- Housing cover; 8- Torque measuring device; Fixed Body 9 8 7 6 5 4 3 2 1 FIGURE 2 Previous rolling bearing assembly with heating capabilities (only) The current version of the assembly, as shown in Figure 3 also has a stiffness advantage compared with the old design, Figure 2, especially in the case of the thrust roller/ball bearings. With the improved assembly, the \u201c10-Drive spindle\u201d is in full contact with the rotating race of the rolling bearing; furthermore, it is also much thicker. This spreads the axial load over the upper race more evenly, as it would in a real application. For this reason, the material of the rolling bearing driving element is a steel alloy for the thrust roller/ball bearings (51107, 81107 TN). For the other rolling bearings types that can be tested, Table 3, the rolling bearing driving element can be made of an aluminium alloy for improved heat transfer, Figure 3. Figure 4 shows an exploded 3D view of the previous and current assemblies. This exploded view also shows the assembly order of both assemblies. Despite having more complex parts, the proposed solution is as difficult as the previous one to assemble. At this point, it should be apparent that in Figures 2 and 4A, the assembly is shown with a thrust roller bearing (81107 TN), but in Figures 3 and 4B the new assembly is shown with a tapered roller bearing (32008 X/Q). In fact, with the proposed assembly rolling bearings with dimensions up to the dimensions of the 32008 X/Q tapered roller bearing can be tested",
            " The free volume in the rolling bearing is obtained by comparing the total volume of the rolling bearings to a solid cylindrical ring of the same dimensions. Table 5 shows the technical characteristics of the refrigerating and heating circulator that is used to maintain the reference temperature of the system. The thermal fluid that is used is a mixture of distilled water with 30% of automotive antifreeze fluid. This particular equipment allows an external temperature probe (PT100) as the control point. In the current setup, three different measurement points can be set for temperature control: Ttop, Tsump and Tint as represented in Figure 3. In the case of a grease lubricated test using a 7206 BEP rolling bearing the temperature of the outer race is monitored through a hole in the rolling bearing bracket using Tint (Figures 3 and 5D). After the temperature location for control is selected, a pair of Ttop, Tsump and Tint can be selected for further temperature monitoring using a type K thermocouple. The room temperature is also monitored using a type K thermocouple. These measurements are continuously recorded. The cooling-heating system has an inherently high thermal inertia; this means that once the operating temperature is attained it will be stably kept",
            " The thrust roller bearing 81107 TN was selected; because of all the possibilities (Table 1) the 81107 TN is the least stable. It is the least stable simply due to the fact that the rolling elements are not guided by any of the rings, thus representing a worst case scenario in what regards SD in the frictional torque measurements. Table 7 shows the selected operating conditions and oil properties. The oil kinematic viscosity was measured according to the Engler method IP 212/92.22 The oil sump temperature, Tsump was set to 80 C as shown in Table 7. Figure 3 shows the position and approximate location of the PT100 used to control Tsump. The variability in the manufacturing process, including surface finish, guarantees that any differences that exist between different rolling bearings of the same model are not sufficient to yield significant differences between the measured torque loss at the same operating conditions. These differences should be further mitigated if a running in procedure is conducted. The measurements were performed in 12 different sessions split in half between two consecutive days",
            " This is due to the fact that when the driving motor was turned off for the torque loss measurement, the oil sump started to cool down, the controller detected this temperature deviation from the reference and started increasing the temperature of the cooling-heating fluid, resulting in a slight increase of the temperature at the top of the rolling bearing. It should be noted that Ttop is measured near the heat transfer interface between the drive spindle and the roller bearing inner race-driving element, Figure 3. Figures 7 and 8 show the results presented in Tables 8 and 9. In Figure 8, the bandwidths associated to the experimental results represent the average value of the measurement \u00b1 the SD (Tables 8 and 9). One of the advantages of performing a rolling bearing torque loss test in conditions that go from boundary Speed (rpm) 100 200 400 800 1200 1600 Tsump ( C) 79.4 \u00b1 0.1 79.1 \u00b1 0.1 79.2 \u00b1 0.2 79.3 \u00b1 0.1 79.4 \u00b1 0.1 79.6 \u00b1 0.2 through full film lubrication is that the reference values for \u03bcbl and \u03bcEHL, 7 for a particular rolling bearing geometry and lubricant, can be properly found by optimization as described in Reference 19"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000478_s12206-015-0703-z-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000478_s12206-015-0703-z-Figure5-1.png",
        "caption": "Fig. 5. Pad moment balance geometrical relation.",
        "texts": [
            " Fluid momentum acting on the pad can be omitted due to the small order compared with pressure moment. The net pad moments, by summing pressure moment of all elements NE, should be zero: 0 NE i i d M =\u00e5 uuur . (11) Using surface integration on each pad, the fluid film moment differential about the pad pivot is found by taking the cross product of vector ir ur with the differential force vec- tor id F uur , i.e. i i id M r d F= \u00b4 uuur ur uur . (12) Neglecting the pad thickness\uff0cthe vector ir ur in Fig. 5 can be obtained, ( ) ( )1 cos sini pr R Rg x g h\u00e9 \u00f9= - +\u00eb \u00fb r r r . (13) The differential of the fluid film force vector can be written as: i i id F dF dFx hx h= + uur r r , (14) where cosi idF P Rd dzx g g=- and sini idF P Rd dzh g g= . The differential moment is 2 sini id M PR d dzg g k= - uuur ur . (15) k x h= \u00b4 ur r r . (16) The acting force on the pad will transmit to the compliant shim and Fs i = Fp i. With the variable stiffness shim, the stiffness of the shim can be adjusted according to the required working states"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003449_j.measurement.2020.108458-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003449_j.measurement.2020.108458-Figure1-1.png",
        "caption": "Fig. 1. Three-dimensional view of a shoe in the sagittal plane at an inclination angle of (\u03b1) showing distances measured by the infra-red sensors S1 and S2.",
        "texts": [
            " A pair of infra-red distance sensors were used to measure FGA based on simple two-dimensional geometry. The IR sensor (from Sharp Corporation, Japan) consists of IR light-emitting diode and a photodiode that works as a transmitter and detector, respectively. The sensor works on the principle of triangularization [30] and provides a voltage (in volts) as the output that depends on the distance of the object from the sensor. The sensor provides a line of sight distance with a minimal beam spread at a longer distance, which ensures an accurate measurement of the far distant object. As shown in Fig. 1, \u03b1 is the FGA in the sagittal plane, which can be measured using a pair of IR sensor S1 and S2 fixed at a distance L apart on the shoe. The distance d1, d2 are the line of sight distances from the sensors S1 and S2 whereasd1Xcos(\u03b1), d2Xcos(\u03b1) are the perpendicular distances from the sensors. As per the (d1 \u2212 d2)Xcos(\u03b1) = LXsin(\u03b1) (1) FGA(\u03b1) = tan\u2212 1(d1 \u2212 d2) L (2) geometric calculation, \u03b1 in the sagittal plane is given by the equation (2). The FGA was validated against the gold standard 3D motion capture system (from BTS Bioengineering, USA) in our previous study [31]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure11-1.png",
        "caption": "Fig. 11. The CHTC of the recirculating hollow-shaft with cone angle at an inlet flow rate of 3L/min and a rotational speed of 3000rpm.",
        "texts": [],
        "surrounding_texts": [
            "The internal wall surface friction loss is mainly judged by the wall shear stress obtained by CFD. 3514 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. Fig. 8 and Fig. 9 illustrate the friction loss increases with both rotational speed and flow rate. This is because the increase in rotational speed and flow rate exacerbates the irregular pulsation of the fluid. It is noted that friction loss of recirculating hollow-shaft is higher than direct-through hollow-shaft. III. STRUCTURE OPTIMIZATION OF RECIRCULATING HOLLOW-SHAFT From the analysis of the CHTC above, it is noted that the CHTC in the middle region is small and uniform compared to bottom wall and top wall of the recirculating hollow-shaft. In Section II, the phenomenon of a small CHTC at the bottom of the recirculating hollow shaft is analyzed when the rotational speed is 3000 rpm. In order to solve this problem, the bottom tail baffle shown in Fig. 3 is changed into a coneshape. Fig. 10 shows the velocity contour and vector diagram of the recirculating hollow-shaft with cone shape. Comparing Fig. 10 and Fig. 6(b), it can be seen that the cone angle facilitates the formation of a stronger toroidal vortex flow in the bottom area. As shown in Fig. 5(b), the CHTC at the bottom of the original structure is around 250 W/m2/K. However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization. Although changing the bottom to a cone-shape can improve the CHTC of the bottom area, the overall CHTC does not increase much. A simple optimization method is to gradually reduce the inner diameter of the bottom stationary cooling tube, so that it enhances the disorder of the bottom fluid, thereby improving the value of the overall average CHTC. The detailed structure is depicted in Fig. 12. Fig. 13 shows that the maximum value of the CHTC of the structure after optimization is 1.57 times that before optimization, and the average CHTC is increased by 16%. However, the friction loss of the wall surface increased by 78%. In addition, the pressure drop of the system is greatly increased due to the reduction of the inner diameter of tube. Obviously, from the overall consideration, reducing the inner diameter of tube is not a good optimization method. 3515 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. A recirculating hollow-shaft structure with blades at the bottom is proposed in this paper. Its specific structure is shown in Fig. 14. There are 6 blades distributed equidistantly in the circumferential direction in the bottom area. By comparing Fig. 15 and Fig. 6(b), it can be found that the flow of the recirculating hollow-shaft structure in the blade area is more turbulent than before, which will greatly increase the CHTC on the surface of the blade area. The results shown in Fig. 16 well verify the above analysis. The maximum CHTC at the corner of the blade, the local CHTC of the blade area can reach 2038W/m2/K, and the average CHTC of the entire heat exchange surface is 670W/m2/K, which is 49% higher than before optimization. Although the blade can increase the CHTC at the bottom, it also increases friction loss. The pressure drop between the inlet and outlet of the optimized structure increased by 2.3% compared with the original structure. In the case of a rotational speed of 3000rpm and a flow rate of 3 L/min, the comparison of specific data before and after optimization is shown in Table II. The rotor experimental platform shown in Fig. 2 uses eight heating pipes evenly distributed on the outer edge of the rotor, which is used to equivalent the actual rotor loss. In order to highlight the optimization of the cooling effect of the recirculating hollow-shaft, the rotor experimental platform was modeled and the temperature field simulation was performed using Workbench software. It can be seen from Fig. 17 that the rotor average temperature of the optimized structure is lower than that of the original structure, and the temperature difference between the two increases as the rotor loss increases. Fig. 18 shows that at a rotor loss of 250W, the temperature of the optimized structure is 33.6oC lower than the original structure. 3516 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSIONS This paper has evaluated and compared the CHTC and friction loss of recirculating hollow-shaft and direct-through hollow-shaft. It shows that the recirculating hollow-shaft is superior to the direct-through hollow-shaft in terms of the CHTC, but friction loss of direct-through hollow-shaft is lower. The tendency of the CHTC of the recirculating hollow-shaft to be higher than that of the direct-through structure increases with rotational speed, but decreases with the increase of flow rate. When flow rate is low and rotational speed is high, the recirculating hollow-shaft has obvious advantages. Furthermore, the optimization of the recirculating hollow-shaft structure has been done in this paper. It is found that the blade and cone angle have significant influence on the CHTC. The optimized structure of changing the bottom tail baffle into a cone-shape and adding blades to the bottom can increase the CHTC by 49%. Compared with the original structure, the optimized design structure can reduce the rotor temperature by 33.6\u00b0C when the rotor loss is 250W. It has the advantage of efficiently cooling the rotor of the high speed permanent magnet synchronous machines. This paper focuses on simulation analysis, and specific experiments will be conducted in the future to verify the simulation results."
        ]
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure15.6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure15.6-1.png",
        "caption": "Fig. 15.6 Hydrodynamic lubricated a plane bearing and b textured bearing",
        "texts": [
            " (iv) Partial texturing shows better result for friction reduction than full texturing. Recently, Bifeng et al. (2018) have proposed a novel texturing scheme on piston ring bymodifying barrel-shaped ring scheme in the formof variable texture depth. They found that average friction power for novel texture pattern is 10.04% lesser than normal texturing. While, in comparison with barrel shape, it has obtained 16.85% lesser average friction power. The plane and textured hydrodynamic journal bearing as shown in Fig. 15.6 is well established. Although, limited research is carried out on engine bearings with texture effect. Surface texture adopted in engine bearings can decrease the friction coefficient, noise intensity, fuel consumption and increase in load carrying capacity. This ultimately reduces energy consumption to cause a decrease in fuel consumption and engine emissions (Ligier and Noel 2015). In modern times, surface texture analysis of tires in automobiles can be studied to control the friction for safe driving at various road conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000063_wemdcd.2019.8887847-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000063_wemdcd.2019.8887847-Figure10-1.png",
        "caption": "Fig. 10. Configuration of the machine windings and magnets when a static axis misalignment of 4mm exists. (red color: the coils in which the rms value of the EMF is increased, orange color: the coils in which the rms value of the EMF is decreased, in a cycle: the coils in which the rms value of the EMF remains approximately constant)",
        "texts": [
            " The harmonic of 80 Hz could be used as fault indicative harmonic because it is not related to the magnetic saturation as in this machine are not used ferromagnetic materials. The rest harmonics, that in the faulty case present amplitude variation, will be investigated in further study. In this section the static axis misalignment is studied. Fig. 9 presents the EMF waveforms of the coils of phase A, when a static axis misalignment 4mm exists in the generator. The misalignment was held along the y-axis direction as it is shown in Fig. 10. The coils of decreased EMF amplitude are presented with orange color, the coils with increased EMF amplitude are presented with red color, while the coils in which the EMF amplitude remains roughly unaffected are presented in a cycle in Fig. 10. The fault changes the relative position between coils and magnets which results in some coils the increment of EMF amplitude and in others the reduction. Fig. 11 depicts the variation of EMF rms values of the coils with the variation of axis misalignment. The EMF rms voltage of coils A1, B1, B4, C1 and C4 presents increment when the level of axis misalignment increases. Contrariwise the EMF rms voltage of coils A3, B2, B3, C2 and C3 presents decrement, while the EMF rms voltage of coils A2 and A4 remains approximately constant in relation to the variation of the axis misalignment level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000829_1.4867171-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000829_1.4867171-Figure1-1.png",
        "caption": "FIG. 1. (a) Linked rigid rod model of a flexible fiber with Nseg rigid sphero-cylinders connected by ball and socket joints. ri is the center of mass of segment i, and pi is its unit orientation vector. (b) Bead chain model of a rigid straight fiber (b is the radius of each sphere, L is the length of the fiber, and p is the unit orientation vector).",
        "texts": [
            " The bead chain model treats fibers as rigid chains of spheres, which is useful for simulating rigid fibers, and comparing with results for suspensions of spheres (i.e., when there is only one sphere in the chain). Hydrodynamic interactions are not included in either model. The linked rigid rod model has been employed previously to simulate the motion of flexible, straight, and nonstraight fibers in shear flow.34, 35 Here a fiber is modeled as a chain of Nseg rigid cylindrical segments with hemispherical end-caps connected kinematically with ball and socket joints (Fig. 1(a)). Each segment has half length and radius b. The length of the fiber is L = 2Nseg . The aspect ratio of the fiber is rp = Nseg /b = Nsegrps , where rps = /b is the aspect ratio of one segment. The orientation of segment i is the unit vector pi. This model features realistic fiber properties including fiber flexibility, nonstraight equilibrium shapes, and static friction between segments in contact. The equations of motion for segment i neglecting inertia can be simplified to Fhyd i + Fext i + NCi\u2211 j Fcon i j + Xi+1 \u2212 Xi = 0, (1) Thyd i + Text i + NCi\u2211 j [ Gi j \u00d7 Fcon i j ] + pi \u00d7 [Xi+1 + Xi ] + Yi+1 \u2212 Yi = 0, (2) where Fhyd i , Thyd i are the hydrodynamic drag force and torque, Fext i , Text i are the external force and torque, Fcon i j is the mechanical contact force, Gij is the vector from the center of segment i to the point of contact with segment j, NCi is the number of contacts of segment i, and Xi, Yi are the constraint force and torque that keep the fiber at a constant contour length (i",
            " In this paper we only consider static friction, which is subject to the Coulombic friction law \u2223\u2223\u2223Ffric i j \u2223\u2223\u2223 \u2264 \u03bcstat \u2223\u2223\u2223FN i j \u2223\u2223\u2223, where \u03bcstat is the coefficient of static friction. These friction forces are enforced as constraints; if the calculated friction force for a contact is larger than \u03bcstat \u2223\u2223\u2223FN i j \u2223\u2223\u2223, the contact is broken, and the friction forces are recalculated. More details about the calculation of the friction forces are provided in Ref. 39. In the bead chain model illustrated in Fig. 1(b), each fiber is treated as a straight chain of Nk osculating spheres of radius b (diameter d = 2b). This model is employed so that suspensions of This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 193.6.61.40 On: Thu, 03 Apr 2014 14:34:06 spheres (Nk = 1) and suspensions of straight fibers (Nk > 1) can be investigated using the same model features and assumptions; we do not employ this model to investigate suspensions of curved fibers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002061_1.4033362-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002061_1.4033362-Figure2-1.png",
        "caption": "Fig. 2 Crankshaft-bearing system",
        "texts": [
            " (6) Read the dynamic nodal force on the journal surface of crankshaft from the data file, and apply them on the journal surface of crankshaft. Then, carry out the finite element analysis of crankshaft by ANSYS software to obtain the dynamic stress on the surface of crankshaft. (7) Calculate the fatigue strength of crankshaft. The flowchart of the calculation about the strap-on type analytical method is shown in Fig. 1. 2.2 Analytical Model 2.2.1 Dynamical Model. The crankshaft-bearing system of a four-cylinder engine is shown in Fig. 2. When the coupling effect of the tribology of main bearing is considered, the constraint of main bearing housing is taken out and replaced with the resultant oil film force of main bearing. The resultant oil film force of main bearing is obtained by solving the Reynolds equation. The finite element model of crankshaft is established by the three-node quadratic Timoshenko beam element in ANSYS software [10], in which the journal of crankshaft is simulated by the circular cross section element, the crank web is simulated by the equivalent Contributed by the Tribology Division of ASME for publication in the JOURNAL OF TRIBOLOGY",
            "org/about-asme/terms-of-use rectangular cross section element, and the belt pulley and flywheel are included in the simplified form of the equivalent lumped mass element. The finite element model of crankshaft (shown in Fig. 3) is composed of 52 elements and 105 nodes. The information about the geometry, inertia and modal of crankshaft are transformed to the ADAMA software by the modal neutral file (the interface between ANSYS software and ADAMS software) [11]. The dynamic model of crankshaft-bearing system is established by replacing the rigid crankshaft (shown in Fig. 2) with the elastic crankshaft (shown in Fig. 3) in ADAMS software. This model is a mixed dynamic model considering the tribological effect of main bearing. Based on the Lagrange\u2019s dynamical equations, the flexible multibody dynamical governing equation of elastic crankshaft in generalized coordinates is as follows [11]: M\u20acn \u00fe _M _n 1 2 @M @n _n T _n \u00fe Kn\u00fe f g \u00fe D _n \u00fe @W @n T k \u00bc Q (1) where n; _n; and \u20acn are the generalized coordinates and its derivative of elastic crankshaft, M and _M are the mass matrix and its derivative of elastic crankshaft, K is the generalized stiffness matrix, fg is the generalized gravitational force, D is the damping matrix, W is the constraint matrix, k is the Lagrange multiplier, and Q is the generalized force matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001533_bit.24905-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001533_bit.24905-Figure1-1.png",
        "caption": "Figure 1. Conceptual representation of the electrochemical cell component of the ESOC.",
        "texts": [
            " The system relies on coulombic titration in order to balance the electrochemical reduction of oxygen with the algal production of oxygen to achieve a steady DO set-point. This method allows one to quantify the algal oxygen production whilst simultaneously reducing oxygen, maintaining a desired DO concentration. This system enables studies similar to that of Raso et al. (2012), whilst allowing in situ, semi-continuous measurement, and DO control without the requirement for gas input thereby allowing one to study either open or closed growth systems. A schematic of the electrochemical system for oxygen control (ESOC) is depicted in Figure 1. The algal chamber is the primary algal growth vessel where algae photosynthesize and generate oxygen. This oxygen is then reduced at the cathode of the electrochemical cell through the addition of electrons (delivered via the electrical circuit) and protons (delivered through the cation exchange membrane). It is this reduction reaction around which the entire system is based for the measurement of oxygen production and its control. A second reaction (oxidation) is required to complete the electrochemical cell, in this case the commonly used 10mM Fe (CN)6 4\u00fe/Fe(CN)6 3\u00fe redox couple was used for its simplicity as an electron donor system (Gerischer, 1969)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000039_aim.2019.8868900-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000039_aim.2019.8868900-Figure1-1.png",
        "caption": "Figure 1. Vertically-articulated SCARA type two-link manipulator affected by external force in the absolute X-Z coordinate system.",
        "texts": [
            " In this paper, an external force estimator based on a RNN trained to acquire an inverse dynamics model of a manipulator is proposed. In the proposed method, the net external force can be calculated from the difference between the joint torque calculated using RNN from the motion of the manipulator and the actually-measured joint torque. In this chapter, the equation of motion of a two-link manipulator used in a simulation to verify the effectiveness of the proposed method is briefly described. As shown in Fig. 1, a selective compliance assembly robot arm (SCARA) vertically articulated is considered. Table I lists the definitions of parameters and variables of the two-link manipulator. The position of the end effector is given by 1 1 2 1 2 1 1 2 1 2 sin sin( ) cos cos( ) l lx l lz                     , (1) and consider the case that external force in the absolute coordinate system  Td x zf ff (2) is applied. Given that the generalized coordinate is  T1 2 \u03b8 and the generalized torque is  T1 2 \u03c4 , the equation of motion is obtained by the Lagrangian method as         2 2 2 1 1 1 1 2 1 2 1 2 2 2 1 2 2 2 1 2 2 2 2 2 2 1 2 2 1 2 2 1 1 1 1 1 2 1 1 2 12 2 2 g g g g g g g g m l J m l l l l C J m l l l C J m l l S c m g l S m g l S l S                                                  (3)    2 2 2 2 2 1 2 2 2 1 2 2 2 2 2 2 1 2 2 1 2 2 2 2 12 g g g g g m l l l C J m l J m l l S c m g l S                                (4) 978-1-7281-2493-3/19/$31"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000622_aim.2015.7222745-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000622_aim.2015.7222745-Figure3-1.png",
        "caption": "Fig. 3. The eneralized search tree in the case of an n degrees of freedom kinematic chain",
        "texts": [
            " These parameters become in geometrical expressions where multiplication with zero eliminates terms of the equation. The appearance of these patterns can be traced back to certain assignment of the DH frames since they depend on the number of 0 values in the DH parameter table. By using the above method it can be sure that all possible combinations are available and the required pattern can be identified in at least one of them. In order to better illustrate the functioning of the presented method the kinematic chain presented in figure 3 is considered. The input to the algorithm are the relative geometrical dimensions, the orientation of the axes in expressed in the global frame and the orientation of the TCP also expressed in the global frame. All these inputs are easily identifiable also by a novice in robotics. In figure 4 the best and the worst from the total of 16 found Tier0 (without dummy frames) solutions are presented for the manipulator presented in figure 3. The runtime of the algorithm was under 1 second. As you can see the two solution differ in the orientation of their frames. The solution considered best has more axes with common orientation and direction. This variation in frame orientation influences also the form of the geometrical relations of the kinematics problem. With the growing popularity of robots beginner roboticists have to be aided by methods and tools that do not require experience in order to be applied or used. Assigning the DH parameters of a robot is a typical example for an activity that requires some experience and is not a straight forward process for one lacking this experience"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000506_978-3-7091-1379-0_2-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000506_978-3-7091-1379-0_2-Figure1-1.png",
        "caption": "Figure 1: Kinematic scheme of proposed mechanism.",
        "texts": [
            " Singularity analysis is based on both Jacobian matrix (Gosselin and Angeles, 1990) and screw theory (Dimentberg, 1965, Glazunov, 2010) and dynamics is analyzed by LagrangeD\u2019Alembert principle. All obtained theoretical results are tested on a virtual model of the mechanism within MATLAB/Simulink environment. The main contribution of this paper is that the new type of the mechanism is discussed and analyzed. It is also shown that the proposed mechanism has no singularities within the workspace. The proposed mechanism is shown in Figure 1. Each leg of the mechanism is constructed as follows: \u2013 the axis of the first revolute joint of i-th (i-1, 2, 3) leg is x-, y- or z- axis (for Leg 1, Leg 2 and Leg 3, respectively) of the three dimensional Cartesian coordinate system; \u2013 the axes of the second and the third revolute joints in each leg are orthogonal to the axis of the first revolute joint in the same leg and are parallel to each other; \u2013 axes of the fourth and the fifth revolute joints are parallel to the axis of the first revolute joint. Note that all three legs are symmetrical and the first R-joint in each leg is actuated. One can see that in the initial configuration of the mechanism (as shown in Figure 1) all the angles between links in each leg are right angles and following conditions must be satisfied: l4 = l2; lA + l2 = l1 + l3 + l5. (1) Here l1 = AiBi, l2 = BiCi, l3 = CiDi, l4 = DiEi, l5 = EiF, lA = OAi. Taking (1) into account, it was found that for the discussed mechanism a basic system of equations that represents a relationship between Cartesian coordinates x, y, z and generalized coordinates \u03b81, \u03b82, \u03b83 can be written as follows:   (y \u2212 lx sin \u03b81) 2 + (z + lx cos \u03b81) 2 \u2212 l22 = 0; (z \u2212 ly sin \u03b82) 2 + (x+ ly cos \u03b82) 2 \u2212 l22 = 0; (x\u2212 lz sin \u03b83) 2 + (y + lz cos \u03b83) 2 \u2212 l22 = 0; (2) where lx = l2 \u221a 1\u2212 ( x+ l2 l2 )2 , ly = l2 \u221a 1\u2212 ( y + l2 l2 )2 , lz = l2 \u221a 1\u2212 ( z + l2 l2 )2 This system of equations can be used to solve forward and inverse kinematic problem, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003004_pesgre45664.2020.9070363-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003004_pesgre45664.2020.9070363-Figure4-1.png",
        "caption": "Fig. 4. Meshed FE model of a BLDC motor showing uniformly demagnetized magnetic flux density (BM) profile.",
        "texts": [
            " The change in electromagnetic quantities, which are also used as fault signatures shall be taken into consideration for SITF and demagnetization fault diagnosis. Demagnetization fault is emulated in the BLDC motor by changing the magnetic coercivity HC value of the PMs. In this study, the coercivity value of two adjacent poles of the rotor has been changed by 40%, in order to induce uniform demagnetization effect in the PMs of the machine. The demagnetized model with the flux density profile is shown in Fig. 4. The change in machine quantities during the demagnetization fault effect can be better understood through the Magnetic Equivalent Circuit (MEC) Analysis. Magnetic Equivalent Circuit (MEC) Analysis. It has been elaborated earlier in [9-10] regarding the direct relation of change in magnetic quantity of a PM to that of magnetic flux density (BM) and further onto the back-EMF (EB) and flux linkages ( ) produced by the PM in a BLDC motor. This can be validated from (14)-(16). The magnetic equivalent circuit of a PM along with its B-H characteristics (at 80\u00b0C used in this study) can be given from Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002109_j.proeng.2015.12.162-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002109_j.proeng.2015.12.162-Figure2-1.png",
        "caption": "Fig. 2. The experimental unit: 1 \u2013 one-component accelerometers x10; 2 \u2013 flexible ropes x4",
        "texts": [
            " Natural frequencies of bending vibrations of the rotor in the range from 0 to 120,000 rpm (2,000 Hz) are represented in table 1, frequencies and shapes of torsional and longitudinal vibrations are not considered. As it is seen from the table 1, frequencies corresponding to the third and fourth bending shape lie in the prohibited range 45,500\u201384,500 rpm. In accordance with the results of the experiment the equivalent model of the rotor of the starter-generator is built with natural frequencies and shapes close to the results of the experiment. The model is built by criteria of equality of masses, lengths and first natural bending frequencies of an equivalent model and its real prototype (Fig. 2) [13, 14, 15]. The results of calculation of natural frequencies on the equivalent model of the SGR are represented in table 3. Thus, the equivalent model of the SGR is built where both first and second natural frequency and shape of bending vibrations. A simplified beam FEM model of the rotor (Fig.3) is arranged in the package Ansys Mechanical APDL, and calculation of critical speeds of a rotor [4] is performed in a wide range of stiffness of the assembly [16]. As calculation shows (Fig. 4), at bearings stiffness less than 106 N/m, flexible assemblies have almost no influence on critical speeds and shapes of the rotor, so it can be considered on free-free conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002341_ijrapidm.2016.078746-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002341_ijrapidm.2016.078746-Figure4-1.png",
        "caption": "Figure 4 (a) Circularity tolerance; (b) CMM measurement",
        "texts": [
            " Dimensional accuracy Standard cube model having the simplest geometry with overall dimensions of 30  30  30 mm3 (see Figure 3a and 3b) and hole diameter of 10 mm printed on the ELAM R3D2 machine to understand its resolution and accuracy is used. The cube was printed six times to examine the repeatable accuracy of the ELAM. The objective was to investigate the dimensional accuracy which includes linear dimensions along x-, y- and z-axis and diameter as well as the circularity (Figure 3a) of the hole in the cube. Digital Vernier calliper with least count 0.02 was used for measurement of linear dimensions. For diameter and circularity, touch probe Coordinate Measuring Machine (CMM) was employed as shown in Figure 4(b) to have a greater level of accuracy in measurement. The results of linear dimension, diameter and circularity measurement for all cubes are reported in Table 3. Figure 5 clearly shows the error distribution along different axes. It was found that during printing, the dimensions along x and y directions are undersized as clearly seen in Figures 5(a) and (b); all the error bars are below the nominal dimension (30 mm). Further, the dimension along the z-direction was oversized as seen in Figure 5(c)",
            " The possible reason was high perimeter width setting in the Slic3r software, temperature difference on the bed surface and changes in environmental conditions. The flow rate setting for the perimeter was reduced in the slicing software and uniform temperature of the bed surface was maintained. Also, the environmental conditions like humidity and temperature rise were maintained at suitable value. These factors can only minimise the dimensional inaccuracy as it is an inherent problem in the AM machines due to the approximations in STL file. Circularity According to ISO/DIS 1101-1996, the circularity tolerance (Figure 4a) is the minimum radial distance between two concentric circles enclosing the given feature (ISO, 1996). For defining a hole, measurement of six points on its circumference was performed using computer-controlled touch probe CMM. The bar chart as depicted in Figures 5(d) and (e) indicate three issues: 1 The circularity of six circles measured at the top and bottom surface. 2 The difference between measured diameter and nominal diameter as diameter error. 3 The diameter error is shown as a percentage of the nominal diameter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001836_iecon.2015.7392422-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001836_iecon.2015.7392422-Figure5-1.png",
        "caption": "Fig. 5. Kinematics of the robot arm.",
        "texts": [
            " er and \u03b8r are cross product and angle of the unit vecotr (e0 = [ 1\u221a 3 , 1\u221a 3 , 1\u221a 3 ]T ) and the vector from the destination r0 to end effector r (r \u2212 r0). er = e0 \u00d7 (r \u2212 r0) |r \u2212 r0| (7) \u03b8r = atan2 (|e0 \u00d7 (r \u2212 r0)|, e0 \u00b7 (r \u2212 r0)) (8) The distances d1, d3, d3 from end effector to each perpendicular surfaces can be calculated as (9). dn = |nn \u00b7 (r \u2212 r0)| (9) The L-J potential field can be expressed by the sum of the influence in each surface as (10), and the force can be calculate form the potential field by (11) 002167 U = U(d1) + U(d2) + U(d3) (10) F (r) = \u2212\u2207U (11) The kinematics of robot arm is shown in Fig. 5. The joint parameters is shown in Table I. The direct kinematics can be calculated from the composition of homogeneous transformation form joint space \u03a3i\u22121 to joint space \u03a3i as (12). 0T5 = 0T1 1T2 2T3 3T4 4T5 (12) Here, the homogeneous transformation form joint space \u03a3i\u22121 to joint space \u03a3i was defined by joint parameters and (30) i\u22121Ti = \u23a1 \u23a2\u23a3 Ci \u2212Si 0 ai\u22121 C\u03b1i\u22121 Si C\u03b1i\u22121 Ci \u2212S\u03b1i\u22121 \u2212S\u03b1i\u22121 di S\u03b1i\u22121 Si S\u03b1i\u22121 Ci C\u03b1i\u22121 \u2212C\u03b1i\u22121 di 0 0 0 1 \u23a4 \u23a5\u23a6 (13) Here, the Sn and Cn represent sin(\u03b8n) and cos(\u03b8n). The meaning of homogeneous transformation in (30) is shown in (14)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003700_adem.202000958-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003700_adem.202000958-Figure1-1.png",
        "caption": "Figure. 1. (a) Schematic of the UAM system; (b) Schematic diagram of UAM sample, and an image",
        "texts": [
            " All rights reserved manufacturing (VHP UAM). Three typical processing conditions were employed, which represented different levels of input energy. The microstructure and texture evolution among layers have been studied by the EBSD technique, with special focus on the dynamic recrystallization (DRX) and dynamic recovery (DRV) at the plastic deformation affected and unaffected regions. In this study, the tested samples were fabricated with a 9kW ultrasonic additive manufacturing machine developed at Harbin Engineering University. Figure. 1(a) shows a schematic of the UAM system. The commercial Al-1100 tapes with a thickness of 200\u03bcm and a width of 24.5mm were consolidated on Al-6061 substrates (170mm (length)\u00d7110mm(width)). Weld speed was fixed at 25mm/s and the process was conducted at room temperature. During the UAM process, the metal tapes were manually laid on the working platform. The clamping and pre-tension system was realized by pneumatic cylinders and the air pressure was set at 0.4 MPa. Samples with five layers of Al-1100 tape were prepared for each processing condition. Figure. 1(b) shows the schematic diagram and image of a five-layer ultrasonically consolidated sample. To investigate the effect of the various processing conditions and ultrasonic energy levels on the microstructure evolution in the built components, three processing parameter combinations were used here, they are the low amplitude (25\u03bcm)/low normal force (1000N) condition (named as Case I), high amplitude (35\u03bcm) /low normal force (1000N) condition (named as Case II), and high amplitude (35\u03bcm)/high normal force (1600N) condition(named as Case III), respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003711_tmrb.2020.3042992-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003711_tmrb.2020.3042992-Figure2-1.png",
        "caption": "Fig. 2. Achieving a haptically continuous proxy PSM update: (a) proxy and real PSM discrepancy before the update, (b) proxy PSM position update.",
        "texts": [
            " For example, if the interaction force between the proxy PSM and the model environment is larger than that of the PSM on the real environment, the position update will place the proxy PSM higher than it was before the update. The user, trying to regulate forces, will feel a decrease in the force feedback, and will push down again. However, the next update will again replace the proxy PSM higher, effectively resulting in users feeling like they can \u201cfall through\u201d the model environment. Therefore, a simple continuous update should be avoided. A more complete update strategy is shown in Fig. 2 which allows model interaction while also preserving haptic continuity, i.e., no jumps in the MTM interaction forces during nominal operation. In Fig. 2a, the current position of the PSM, its local surface normal and closest point to the model surface are represented as p, n\u0302, and s, respectively. The current position of the proxy PSM, p\u0303i corresponds to a closest point s\u0303i on the model surface, with a penetration depth di along the local surface normal \u02dcn\u0302i. In Fig. 2b, the updated proxy PSM position, p\u0303i+1, is given by: p\u0303i+1 = s \u2212 di+1n\u0302 (3) Assuming that location s\u0303i has an associated normal stiffness ki then the magnitude of the force calculated by the model is fi = kidi. After the proxy update, the magnitude of this force should be constant, i.e., fi+1 = fi. Therefore, assuming the normal stiffness at location s is ki+1, then the penetration depth, di+1, should be: di+1 = diki ki+1 (4) In in this study, the environment had a constant stiffness, so di+1 = di"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000256_icar46387.2019.8981551-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000256_icar46387.2019.8981551-Figure4-1.png",
        "caption": "Fig. 4. Empirical joint pdf of P ,\u03b8 and P ,\u03c6 and fitted pdf for individual scenarios. (a) OLoS corridor. (b) NLoS corridor. (c) Students laboratory. (d) Computer room. (e) Cafeteria hall.",
        "texts": [
            " In order to evaluate whether the selected pdf is consistent with the empirical pdf, the Kolmogorov\u2013Smirnov testing is further used to judge whether the null hypothesis that both pdfs are drawn from the same distributions is true. Without specifically mentioning, the analytical pdfs presented in the following results have passed the Kolmogorov\u2013Smirnov testing and proved to fit well with their empirical counterparts. 1) Statistics of P ,\u03b8 and P ,\u03c6: As shown in the model (1) of the received signal contributed by an MPC, the transmitted waves with \u03b8- and \u03c6-polarization propagate along the same route for a fixed path. Thus, it is natural to consider that P ,\u03b8 and P ,\u03c6 are correlated for the same path. Fig. 4(a)\u2013(e) depicts, respectively, the empirical joint pdf f(P ,\u03b8, P ,\u03c6) of P ,\u03b8 and P ,\u03c6 for the five scenarios. The bivariate log-normal pdfs fitted to the empirical pdfs are also illustrated. We choose the lognormal distribution for fitting because the power of the four entries in A have been found to follow log-normal distributions. This leads to the conjecture that P ,\u03b8 and P ,\u03c6 are also log-normal-distributed [40], [41]. Table II reports the parameters of the log-normal pdfs fitted to the empirical pdfs. It is worth mentioning that the value of total path power, i.e., P ,\u03b8 + P ,\u03c6 can be determined according to the existing SCMEs by using intercluster statistics. However, no explicit instructions are given in the standard SCMEs for further splitting the path power among the different polarized components. The pdfs f(P ,\u03b8, P ,\u03c6) presented here may be helpful when allocating the appropriate values of P ,\u03b8 and P ,\u03c6 for the MPCs propagating along the same path. It can be observed from Fig. 4 that except for the corridor scenarios, f(P ,\u03b8, P ,\u03c6) are more or less tilted, implying that P ,\u03b8 and P ,\u03c6 are less-correlated in a closed environment, e.g., with waveguide structures as in the OLoS corridor. Our explanation for this phenomenon is that in the waveguide-alike environment, MPCs are less resolvable. The contribution of noncoherent diffuse scattering to each cluster is more significant than in other cases. Consequently, P ,\u03b8 and P ,\u03c6 become less correlated due to severe distortions of the polarizations in such cases",
            "69, whereas the cross-correlation among other ellipse parameters is usually less than 0.1. The scenarios where \u03c1P ,\u03b8,P ,\u03c6 is sorted in the descending order are the hall [i.e., the scenario (e), with \u03c1P ,\u03b8,P ,\u03c6 = 0.69], the NLoS corridor [the scenario (b), with \u03c1P ,\u03b8,P ,\u03c6 = 0.63], the laboratory (the scenario (c), with \u03c1P ,\u03b8,P ,\u03c6 = 0.44), the computer room [the scenario (d), with \u03c1P ,\u03b8,P ,\u03c6 = 0.19], and the OLoS corridor [the scenario (a), \u03c1P ,\u03b8,P ,\u03c6 = 0.11]. Such a sequence is also consistent with the extent to which the pdf of P ,\u03b8 and P ,\u03c6 is tilted, as illustrated in Fig. 4. Based on these observations, it is necessary to define the cross-correlation between P ,\u03b8 and P ,\u03c6 depending on the type of indoor environments, and generate the random realizations of P ,\u03b8 and P ,\u03c6 by using their joint lognormal pdf with particular \u03c1P ,\u03b8,P ,\u03c6 . The other parameters in \u03b3 may be generated independently by using the analytical pdfs consistent with the empirical samples in specific propagation scenarios. IV. MODEL VALIDATION To understand the effect of applying the newly proposed ellipse-parameter-based modeling method, we select a propagation scenario, e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000063_wemdcd.2019.8887847-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000063_wemdcd.2019.8887847-Figure4-1.png",
        "caption": "Fig. 4. Schematic representation of the static axis misalignment [14].",
        "texts": [
            " This condition creates a not uniform air gap in which during the rotor rotation the positions of the minimum and the maximum air gap lengths are not fixed, but their values are constant. In other words, the air gap length varies as a function of the time. Finally, mixed angular misalignment occurs when the rotor symmetrical axis does not coincide with the rotational axis neither with the stator symmetrical axis. Consequently, the air gap is not uniform during the rotation, the positions of the minimum and the maximum air gap lengths do not remained constant, while their values also change [4]-[5]. In Fig. 4 the schematic representation of the static axis misalignment in an AFPM machine is presented. At first the static angular misalignment will be studied for two different levels, 20% and 40%. The rotor displacement leads to a minimum airgap length above the stator coil A2 and a maximum airgap length above the stator coil A4. Fig. 5 depicts the EMF of the coils of phase A for the healthy case and the case that a static angular misalignment 40% exists. The coils A1 and A3 which are aligned with y-axis do not present variation in their EMF amplitude when the fault exists"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000778_20140824-6-za-1003.02511-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000778_20140824-6-za-1003.02511-Figure3-1.png",
        "caption": "Fig. 3. 1-DOF robot arm.",
        "texts": [
            " The update rules for the actor and process model remain the same as in Section 2.2: \u03b8ak+1 = \u03b8ak + \u03b1a\u2207xV (xk+1 \u2212 xr,xr)\u2207uf\u0302(xk,uk)\u2207\u03b8a\u03c0(x) and \u03b8pk+1 = \u03b8pk + \u03b1p(xk+1 \u2212 x\u0302k+1)\u2207\u03b8p f\u0302(xk,uk) The entire model-learning actor-critic (MLAC) algorithm for nonlinear disturbance rejection is summarized in Algorithm 1. Note that other RL algorithms, such as the standard actor-critic, can be applied in this setting as well. The proposed control scheme is applied a 1-DOF robot arm operating under the influence of gravity as shown in Fig. 3. The equation of motion is: Algorithm 1 Nonlinear disturbance rejection via MLAC with fuzzy approximators Input: \u03b3, \u03bb, and learning rates \u03b1 1: Initialize \u03b6j,0,\u2200j, and fuzzy function approximators 2: Apply un,0 + \u2206u0 3: k \u2190 0 4: loop 5: Measure xk+1, rj,k+1,\u2200j and xr 6: Choose un,k+1 according to nominal control law 7: % Choose compensation action and update actor 8: uk+1 \u2190 output of actor rule base 9: \u03b8ak+1 \u2190 \u03b8ak+ \u03b1a\u2207xV (xk+1 \u2212 xr,xr)\u2207uf\u0302(xk,uk)\u2207\u03b8a\u03c0(x) 10: % Update process model 11: \u03b8pk+1 \u2190 \u03b8pk + \u03b1p(xk+1 \u2212 x\u0302k+1)\u2207\u03b8p f\u0302(xk,uk) 12: % Update critic 13: for \u2200j \u2208 [1, 2, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.24-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.24-1.png",
        "caption": "Fig. 18.24 Concept of quadrupole mass spectrometer (a) and miniature QMS assembly made with MEMS technology (b) (from Imperial College London)",
        "texts": [
            " Mass spectrometry is used to determine a specific molecular characteristic called mass-to-charge (m/z) ratio. The spectrometer is fast, has a very wide dynamic range, and is quite efficient, albeit it is also quite expensive. In a quadrupole mass spectrometer (QMS), the separation and determination of the molecule and fragment masses is controlled with high-frequency electric fields produced by the radiofrequency (RF) generator. The RF voltages are applied to four rod-shaped electrically conductive electrodes (quadrupole) placed in high vacuum, as shown in Fig. 18.24a. The filtering and determination of the ionized molecules is then implemented with a programmed modulation of the RF voltage amplitude with simultaneous superimposition of a DC voltage which can also be modulated. The tested molecules are ionized by high-energy electrons, plasma, or with chemical reagents, and enter the space between four rods (poles). The alternating electric field produced by the rods accelerates ions out of the source region and into the quadrupole channel between the rods. As the ions travel through the channel, they are filtered according to their m/z ratio so that only a single m/z value ions can strike the detector inside the ion collector that is positioned at the opposite side from the source",
            " These voltages produce an oscillating electric field that functions as a bandpass filter to transmit only the selected m/z value. Thus, the molecules of the specific m/z are in resonance with the RF frequency and oscillate in the x\u2013y plane, while propagating toward the collector. On the other hand, the out of resonance molecules go astray, hit the rods, and never reach the detector\u2014in other words, they are rejected. The RF and DC fields are scanned (either by DC potential or frequency) to collect a complete mass spectrum. Recently attempts were made to develop QMS in a miniature form by using MEMS technologies, Fig. 18.24b. When the internal energy of a system changes, it is accompanied by an absorption or evolution of heat (as defined by the first law of thermodynamics). Therefore, a chemical reaction, which is associated with heat, can be detected by an appropriate thermal sensor, such as described in Chap. 17. These sensors operate on the basic principles that form the foundation of a microcalorimetry. The operating principle of a thermal sensor is rather simple: a temperature probe is coated with a chemically selective layer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002358_chicc.2016.7553174-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002358_chicc.2016.7553174-Figure1-1.png",
        "caption": "Fig. 1 Fully-submerged hydrofoil craft",
        "texts": [
            " A 2 degree-of-freedom steering model of the fully submerged hydrofoil including roll and yaw dynamics is established, and the wave disturbance is analyzed based on stochastic wave theory. In Section 3, an EDO based nonsingular terminal sliding mode (EDONTSM) controller is developed for the maneuvering of the FSHC. Simulation is carried out in Section 4 to validate the effectiveness of the proposed control method, followed by Conclusions in Section 5. A typical configuration of a fully-submerged hydrofoil craft is shown in Fig. 1. The T-shaped bow foil is equipped with two controlled flaps, acting together. The aft foil has a pair of central flaps and two pairs of ailerons. Struts of the aft foil are equipped with rudders, which is used for roll and yaw dynamics together with the ailerons. And the bow foil and the central part of the aft foil are for longitudinal motion control. Assuming that the surge speed is controlled by an individual propulsion system, the maneuvering model of a typical marine vehicle is shown as follows: ( ) ,J (1) 0 0( , ) ( , ) ( ) ,c dM C u D u G (2) where = , T is a vector of position and orientation with coordinates in the earth-fixed frame, and = , Tp r is a vector of linear and angular velocities with coordinates in the body-fixed frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003387_fuzz48607.2020.9177781-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003387_fuzz48607.2020.9177781-Figure1-1.png",
        "caption": "Fig. 1: The outline of a cloud -based model for a distributed FLS [1].",
        "texts": [
            " This openness can be accomplished by developing the system to be fully web-based and device self-sufficient, particularly by utilising standard formats for data exchange that are both consistent and readable as realistically as possible. For example, an individual client computer may outline its necessary FLS, an input data can be provided for the specified system by the same or other client computer(s), and lastly the same or other client(s) can repossess the calculated output [1]. The outline of a cloud-based model for a distributed FLS is shown in Fig. 1. Authorized licensed use limited to: Carleton University. Downloaded on October 03,2020 at 13:26:21 UTC from IEEE Xplore. Restrictions apply. A web-based data language for FLS characterizations is the key criterion for implementing such an architecture. The current standard for this purpose is the IEEE-1855 (2016), also known as FML [3], an XML-based mark-up language allowing the human readable and hardware-independent definition of an FLS. FML and FML-compatible pieces of software such as JFML [4] are used as the basic design standard in this study, and the extensibility of this standard is a solution to architecture growth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001756_978-3-319-24502-7_10-Figure10.16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001756_978-3-319-24502-7_10-Figure10.16-1.png",
        "caption": "Fig. 10.16 Squeezing motion of a no-slip sphere towards a near-contact planar wall, driven by the constant force F \u00bc jFjez",
        "texts": [
            " For squeezing motion, however, there is still a O\u00f01= ffiffiffi h p \u00de singularity for two droplets with fully mobile interfaces. It is assumed here that the droplet interfacial tension is so large that the flow-induced deviations from the spherical shape are negligible. For more information about lubrication effects under different surface BCs see [14, 66]. As an illustration of the dynamic effect of lubrication consider a no-slip sphere of radius a approaching a stationary, horizontal no-slip wall in squeezing motion (see Fig. 10.16). The sphere is driven by the constant external force F \u00bc jFjez \u00bc Fh. This is a limiting case of Eq. (10.96), for a = a1 and a2 ! 1 with the lower sphere 2 expanded into a planar wall. Here, h denotes now the distance from the wall to the closest surface point of the sphere. It follows that dh\u00f0t\u00de dt \u00bc Vrel\u00f0t\u00de lt0jFj a h\u00f0t\u00de ; \u00f010:97\u00de with the solution h\u00f0t\u00de h0exp lt0jFj a t : \u00f010:98\u00de Here, h0 is the near-contact starting distance with h0=a 0:01, and lt0 is the single-sphere translational mobility defined in Eq",
            " The exponentially slow vertical approach of the sphere to the wall is a good description in reality only until the gap distance h becomes comparable to the surface roughness of the sphere and plane which in fact leads to plane-sphere contact after a finite time. A finite contact time would be reached also due to the van der Waals attraction force between the planar wall and the sphere which in the distance range where lubrication applies scales as O\u00f01=h2\u00de [68]. The preceding section was devoted to the effects of the solvent flow on the dynamics of suspended particles in an unbounded fluid, except for the discussion in Fig. 10.16 of the near-wall settling of a sphere. Most realistic situations involve the presence of a boundary that may considerably change the flow character, and by reflecting the flow incident upon it, may modify the hydrodynamic interactions between the particles. A pronounced example is the effect of sedimentation, where the backflow of fluid due to the presence of a container bottom, however far distant it may be, cannot be neglected in order to correctly determine the sedimentation velocity of dispersed microparticles [9]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000761_rpj-11-2011-0117-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000761_rpj-11-2011-0117-Figure1-1.png",
        "caption": "Figure 1 Schematic of a first-generation FDM machine",
        "texts": [
            " The main goal of this work is to evaluate a systematic approach for the design of additive manufacturing equipment in which both cost and final accuracy of the machine are considered. As the basis of this study, a layout of a positioning system such as the stereolithography or selective laser sintering equipment (Gibson et al., 2010) was selected. In these systems, the displacement of the tool head is a direct result of the action of mechanical elements and does not to involve mirrors or galvanometers. Figure 1 shows a schematic of the selected concept,which is equivalent to a first generation fuseddeposition modelling (FDM) machine, such as a 3D Modelerw or FDM series (Wang, 2010). It is also interesting tonote that variations of this concept were adapted for simultaneous deposition and polymerisation (SDP), laminated object modeling (LOM) and 3D print technologies. Consequently, this work can also be used for the improvement of other similar technologies (Gibson et al., 2010; Cunico, 2011). For this research, a numerical model that allows the identification of the cost and final error of the equipment The current issue and full text archive of this journal is available at www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003914_gt2015-43940-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003914_gt2015-43940-Figure1-1.png",
        "caption": "Figure 1. CAD model rendering of L-2x test coupon.",
        "texts": [
            " Careful characterization of the channel geometry was vital to the success of this study as flow friction and heat transfer results are strong functions of geometric scales. The following is a description of the test coupons used for this study followed by the methodology and results of characterizing the geometric and roughness parameters of the coupons. Eleven coupons were fabricated for this study, all of which were 25.4 mm in length and width. A summary of these coupons and their design specifications is given in Table 1. An image of the CAD model of the L-2x coupon is shown in Figure 1. Only one coupon (Cyl-Al) was made by conventional machining using 6061 aluminum. This coupon contained 14 round channels running along its length that were each reamed to a diameter of 635 \u03bcm. The Cyl-Al coupon was used to benchmark the experimental facility and data reduction methodology as compared to generally accepted correlations found in the literature. All other coupons in Table 1 were manufactured with rectangular channels using DMLS. Five of the coupons (L-1xCo, L-2x-Co, M-1x-Co, M-2x-Co, S-2x-Co) were manufactured by a commercial company"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure4-1.png",
        "caption": "Figure 4. Terminology for shear. Equal and opposite forces (F ) are applied so as to cause angular distortion \u03b3. Strain is given by the ratio of distortion \u0394Lx to unstressed thickness Ly0. Shear stress is normalized by the area S parallel to the forces.",
        "texts": [
            " The exact value of E, however, may differ in tension and compression, and often the strength differs dramatically. Whereas Young\u2019s modulus applies to both tension and compression, a different relationship is needed for shear. The shear stress, \u03c4, is the force applied, Fs, divided by the area parallel to the force, Ss (\u03c4=Fs/Ss), and shear strain is the ratio of the amount of deflection parallel to the force (\u0394Lx), to the object\u2019s unstressed length perpendicular to the force (Ly0), i.e., tan \u03b3 (figure 4). (At low shear strains, the deflection angle \u03b3 itself, in radians, will be essentially the same as the tan \u03b3, so the angle itself is usually used as shear strain for stiff materials.) The shear modulus, G, takes the same form as Young\u2019s modulus: G . 4t g= ( )/ The dimensions are the same as for tension and compression: \u03c4 and G are force per unit area (so Pa in the SI system), and both tan \u03b3 and \u03b3 (measured in radians) are again ratios and thus dimensionless. Elasticity in materials science means something a bit different from its everyday use"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000723_0954406214541433-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000723_0954406214541433-Figure1-1.png",
        "caption": "Figure 1. Types of contacting filament seal considered in current work. (a) Brush seal (adapted from Jahn,7 with permission from ASME). (b) Leaf seal (adapted from Jahn,7 with permission from ASME). (c) Finger seal (adapted from Proctor et al.,8 with permission from AIAA).",
        "texts": [
            " This is despite new types of Center for Hypersonics, School of Mechanical and Mining Engineering, The University of Queensland, Australia Corresponding author: Ingo HJ Jahn, Center for Hypersonics, School of Mechanical and Mining Engineering, The University of Queensland, Brisbane, QLD 4072, Australia. Email: i.jahn@uq.edu.au contacting filament seals, such as the leaf seal5 and finger seal6 being invented and ongoing research in this field. Schematics of the three most common contacting filament seals are shown in Figure 1. The success to which such seals have been employed has varied significantly.9 The author believes that some of the unsuccessful seal installations are caused by differences between actual seal operation conditions and conditions anticipated during design and that this could have been avoided by designing seals that perform well over a wide range of operating conditions. In the current work, it is explored how general seal properties such as filament mechanical stiffness, blowdown, initial build clearance, and their cross-coupling affect seal performance and performance retention, and it is demonstrated that ideal combinations of seal properties exist, that result in seals, which are less sensitive to operating conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001155_sbr.lars.robocontrol.2014.22-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001155_sbr.lars.robocontrol.2014.22-Figure1-1.png",
        "caption": "Fig. 1. The AR.Drone 2.0 quadrotor and the coordinate systems adopted ({w} and {b} are the global and the body coordinate systems, respectively).",
        "texts": [
            "LARS.Robocontrol.2014.22 118 the state estimation technique adopted to use the sensor data available in a KF implementation. After, Section IV discusses the control method adopted, Section V presents an overview of the system architecture, and Section VI shows and discusses some experimental results. Finally, some important observations are pointed out in Section VII. The experimental platform chosen in this work is the AR.Drone quadrotor, from Parrot Inc, in its version 2.0, which is shown in Figure 1, with the coordinate systems dealt with. It is a rotorcraft aerial vehicle commercialized as a hi-tech toy, originally designed to be controlled through smartphones or tablets, via Wi-Fi network, with specific communication protocols. The AR.Drone is easily purchased in the market at a reduced cost and it is quite easy to buy spare parts to keep it operative, as well as for maintenance. In addition, Parrot provides an open source software development kit (SDK), which makes easier to develop communication and control algorithms for the AR"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003446_0954406220959376-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003446_0954406220959376-Figure2-1.png",
        "caption": "Figure 2. Samples for butt-joint tests manufactured in different directions, left: horizontal, right: vertical.",
        "texts": [
            " In order to investigate the effect of the laser power on microstructural characteristics and on the adhesive joints, samples were built using two different laser sources with a laser power (LP) of 400W and 800W, respectively. The parameters were chosen with respect to a small amount of porosity as well as differences in surface roughness. To determine the porosity, the samples were grinded and polished up to 0.5 mm. In detail, the parameters used are listed in Table 1. Microstructural differences were analysed after polishing with a colloidal silica suspension. Two different sets of samples were produced, oriented vertical and horizontal (cf. Figure 2). The joints were realized by using a two component epoxy resin (2C epoxy, Weicon, Germany). This adhesive has a high temperature resistance from 50 up to \u00fe230 C. It is suitable for vertical application and is spot-weldable in non-cured state. The adhesive properties are listed in Table 2. For post-treatment the SLM and die cast samples were cleaned with isopropyl alcohol in an ultrasonic bath before bonding to ensure a surface which is free of dust and grease. Furthermore, a post-treatment with a 50W laser (cleanlaser, Germany) was conducted to analyse if it improves the adhesive strength",
            " The adhesive behaviour was studied by carrying out lap shear tests according to DIN EN ISO 1465:2009 and butt-joint tests according to DIN EN 15870:2009. The number of repetition of each experiment was three. An overview of the tested combinations is listed in Table 3. The shape of the samples for the lap shear tests was 20 80 3 mm3. The two parts were joined with an overlap length of 7mm and an adhesive nominal layer thickness of 0.5mm (not shown for the sake of brevity). The geometry of the butt-joint samples is shown in Figure 2. The adhesively bonded samples were cured for 24 h at room temperature (RT). All tests were carried out at RT in displacement control with a cross-head speed of 10mm/min. The samples were tested without ageing of the bonded joint and after accelerated ageing using cataplasm (according to DIN EN ISO 9142:2004). The samples were stored 14 days at a temperature of 70 C under 100% relative humidity. The surface roughness is frequently employed as a design criterion for an adhesive joint. A higher surface roughness of the adherent support the spreading of the adhesive on its surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000391_978-3-319-02294-9-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000391_978-3-319-02294-9-Figure12-1.png",
        "caption": "Fig. 12 Clamping force behavior with criteria to not exceed their margins",
        "texts": [
            " The results of the third LCA phase are shown in the form of partial graphs presenting environmental burden by particular machine tool components according to the chosen impact categories [15]. 70 M. Iskandirova et al. Fig. 8 Partial results of indicators for ozone layer depletion CO2 eq. Fig. 9 Partial results of indicators for ozone layer depletion CFC-11 eq. Fig. 10 Partial results of indicators for acidification SO2 eq. Fig. 11 Partial results of indicators for photooxidants creation C2H4 eq. Fig. 12 Partial results of indicators for eutrophication PO4 3- e Fig. 13 Legend to Figures 5 to 13 According to the results of this phase, it is possible to make following conclusion about components the most burdening the environment in chosen impact categories: sled with revolving stand, sled and stand the most contribute to global warming and eutrophication; headstock the most contributes to ozone layer depletion; sled with revolving stand, sled, stand and headstock the most contribute to acidification; platform the most contributes to photooxidants creation",
            " The large torque value is caused by the starten of the drive from zero to maximum velocity. Fig. 9 Drive of the wheels Fig. 10 Movement simulation of the drive of wheels 548 M. Dovica et al. The chosen drive has been used for the simulation of other parts via FEM (finite element method) (Figure 11). Upper position of undercarriage has to be fixed, because the mechanism is not self locking. The weight of the robot may lead to the fall of the lower part of the undercarriage. For this purpose, the locking mechanism was designed. (Figure 12) This mechanism consists of servomechanism that moves the pin into the locking position. This solution is also suitable from the energy point of view. Final design of the robot is shown in the Figure 13. The overall dimensions are 280 x 351 x 175 mm and its weight is 3,3 kg. Clearance of the undercarriage is Robot with Adjustable Undercarriage \u2013 The Design and the Simulation 549 adjustable from 120mm to 150mm. Calculated maximum velocity on the surface without any obstacle and roughs is 0,47 m/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002850_s00170-020-05095-2-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002850_s00170-020-05095-2-Figure5-1.png",
        "caption": "Fig. 5 Thermal deformation",
        "texts": [
            " Using the temperature field, the thermal deformation of the spindle can be analyzed by thermal-structure coupling analysis. Because the middle bearings are fixed, however, the other two sets of bearings can move freely along the axial direction, and the constraints of these three sets of angular contact ball bearings are applied to the finite element model. The contact surfaces of each contact pairs are connected through node couplings. Appling the temperature field to the finite element model, the thermal deformation of the spindle can be simulated using the software of ANSYS. Figure 5 shows the thermal deformation of the spindle. Extracting the analysis results, the thermal elongation of the grinding wheel shaft was only \u2212 0.16 \u03bcm. However, the maximum thermal elongation which occurred in the right side of the spindle was 66.6 \u03bcm. That is to say, the proposed optimization method is useful for reducing the thermal elongation of the spindle. To compare the optimization effect, the finite element analysis of the spindle before optimization is carried out. According to the traditional structure of the spindle, the contact thermal conductivities between bearing and sleeve, sleeve and coupling, and coupling and grinding wheel shaft are 120, 120, and 180 (W/(m2\u00b7K)), respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure1-1.png",
        "caption": "Fig. 1. Mechanically adjusted SFPMM (a) All pole (b) Alternate pole",
        "texts": [
            " Alternate polarity PMs are extended from stator tip to the end resulting in significant flux leakages that ultimately effect electromagnetic performance. Moreover, high PMs usage results in increase of eddy current loss in PMs [5]. Additionally, high PM volume increase machine weight and cost. In order to suppress PM volume, reduce machine cost, reduce machine weight, and flux leakages, an overview is carried out in literature on wide range of SFPMM [6, 7]. Flux leakages is tried to suppress by introducing mechanical adjustor as shown in Fig. 1(a) [8]. Similar concept with alternate mechanical adjustor is analyzed in [9] as shown in Fig. 1(b). whereas comprehensive overview is presented in [10]. This methodology significantly suppresses flux leakages however, this method requires extra mechanical accessories which causes undesirable increase of machine volume and weight. Moreover, PM usage is same which means that due to extra mechanical accessories, mechanical adjustor topologies results in increase of cost. SFPMM exhibits outstanding electromagnetic performance, however increase in PM usage further increase machine cost. To overcome PM usage in SFPMM, an overview is carried out in [6] showing different topology of SFPMM with reduced PM usage whereas E-core SFPMM is introduced in [11] which reduced PM usage to half and term as conventional 12/10 stator slot/rotor pole (12S-10P E-Core SFPMM) as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001316_066101-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001316_066101-Figure6-1.png",
        "caption": "Figure 6. Schematic diagram of compression force.",
        "texts": [
            " (17) By taking the derivative with respect to x d2\u03b8 dx2 = \u2212 Fstrut EI d\u03c9 dx (18) where d\u03c9 dx = sin \u03b8. (19) Combining equations (18) and (19) gives d2\u03b8 dx2 = \u2212 Fstrut EI sin \u03b8. (20) Solving equation (20) and substituting the boundary condition F = FP into it, then \u03c9 = 0. The approximate solution is: \u03c9 = 2 \u221a 2l \u03c0 \u221a Fstrut FP \u2212 1 [ 1\u2212 1 2 ( Fstrut FP \u2212 1 )] . (21) It is known from Euler\u2019s formula that the flexure curve is part of a sine curve. In order to simplify calculations, this paper assumes that the strut after bending is part of a circle. Then we can obtain using the geometry relation in figure 6 that x\u2212 x cos l 2x \u2212 \u03c9 = 0(x \u2265 \u03c9) (22) S = l\u2212 2 \u221a x2 \u2212 (x\u2212 \u03c9)2. (23) 3.2.2.1. Critical pressure of elastic deformation. Figure 7 shows that the regular hexahedral unit structure is elastically deformed before it reaches the proportionality limit. It Laser Phys. 23 (2013) 066101 J Sun et al can be obtained through the geometry relationship during compression that: 1\u03b5 = y \u03c1 (24) 1\u03c3b = ET1\u03b5 = ETy \u03c1 . (25) From relationship between 1\u03c3 and the bending moment M we obtain: M = \u222b A y1\u03c3y dA = ET I \u03c1 . (26) Using the expression of the approximate curvature of Euler\u2019s formula we obtain: d2\u03c5 d2x = \u2212 M ET I "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure4-1.png",
        "caption": "Fig. 4. Planar joint (1-3. Rotary joints, 4. Pins for mounting the hydraulic axes, 5. Main pin (end-effector)).",
        "texts": [
            " This extends the IEHSD control tasks to position/force (position/pressure) control), which is very important in controlling heavy manipulators. Wireless network interface controller (WNIC) extends the IEHSD communication capability in Industry 4.0. To eliminate the rotational movement of the piston rod of the hydraulic actuator around its axis, an additional rotation lockout was used. Fig. 3 shows the rotation lockout and provides a description of the individual elements. Pure translational motion of the mobile platform of the PM is possible after using a specially designed planar joint (Fig. 4). Here, the hydraulic actuator piston rods are mounted on pins in the indi- vidual rotary joints. The moving platform is mounted on the main pin, which is the end-effector (EE). The TPM kinematic model with three 3-RRPRR serial chains is shown in Fig. 5. The TPM design consists of a fixed platform, a mobile platform, twelve passive revolute joints (R), including a planar joint (E) and three active prismatic joints (P). The Chebychev-Gr\u00fcbler-Kutzbach (CGK) formula was used to determine the mobility (m = 3) of the hydraulic TPM [23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003783_indicon49873.2020.9342273-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003783_indicon49873.2020.9342273-Figure1-1.png",
        "caption": "Figure 1 Single link manipulator",
        "texts": [
            " But the continuous-time system was discretized and it was defined by (7) and (8) is quite general and covers most practical situations. for example, in [1] using differential evolution, the above problem, we calculate optimal solution using the performance index with constraints. Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 11:37:24 UTC from IEEE Xplore. Restrictions apply. Consider a manipulator model with a single link, which is illustrated in Fig. 1: 2 2 1 ( ) s in ( ( ) ) ( ) ( ) g v t t t u t l m l m l \u03b8 \u03b8 \u03b8= \u2212 \u2212 +&& & (9) Where\u03b8 is the angular position, m is the mass of the end of the rod element, l is the length of the rod, v is the friction coefficient at the pivot point, and u is the applied torque at the pivot point. The above equation (9) can be rewritten by considering m=2kg, 6v kgms= ,and l=1m. Define 1( )z t \u03b8= , 2 ( )z t \u03b8= & , and then, the manipulator model can be written as We have considered that the manipulator was initially at rest condition i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001131_j.amc.2014.01.138-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001131_j.amc.2014.01.138-Figure4-1.png",
        "caption": "Fig. 4. The ortho\u2013parallel manipulator in configurations corresponding to ideals T23; T66 and T67. The figures in different modes were again drawn with the computer program Alaska.",
        "texts": [],
        "surrounding_texts": [
            "We represent the orientation and the corresponding rotation matrix using Euler parameters. Let a \u00bc \u00f0a0; a1; a2; a3\u00de and define the matrix Rby R\u00f0a\u00de \u00bc 2 a2 0 \u00fe a2 1 1 2 a1a2 a0a3 a1a3 \u00fe a0a2 a1a2 \u00fe a0a3 a2 0 \u00fe a2 2 1 2 a2a3 a0a1 a1a3 a0a2 a2a3 \u00fe a0a1 a2 0 \u00fe a2 3 1 2 0B@ 1CA: If a belongs to the unit sphere a 2 S3 R4; jaj \u00bc 1 then R\u00f0a\u00de is orthogonal with det\u00f0R\u00f0a\u00de\u00de \u00bc 1 and we get a polynomial representations for rotation matrices. Note that R\u00f0a\u00de \u00bc R\u00f0 a\u00de and from this it follows that sometimes we may have 2 or more distinct subvarieties of the configuration space which nevertheless correspond to the same physical positions of the mechanism. Perhaps a simple example clarifies the situation. Consider a single rigid body with a constraint 2a2 3 1 \u00bc 0. In this situation I \u00bc hjaj2 1;2a2 3 1i \u00bc I1 \\ I2 \u00bc h2a2 0 \u00fe 2a2 1 \u00fe 2a2 2 1; a3 1= ffiffiffi 2 p i \\ h2a2 0 \u00fe 2a2 1 \u00fe 2a2 2 1; a3 \u00fe 1= ffiffiffi 2 p i: Let us further denote V \u00bc V\u00f0I\u00de and Vj \u00bc V\u00f0Ij\u00de. Now interpreting R as a map R : S3 ! SO\u00f03\u00de then it is clear that R\u00f0V\u00de \u00bc R\u00f0V1\u00de \u00bc R\u00f0V2\u00de; although Vj V and moreover V1 \\ V2 \u00bc ;. From the point of view of analysis of the configuration space it is evident that we can simply take one of the Vj and ignore the other. Since this situation occurs frequently it is convenient to have a notation for it. Suppose we have m rigid bodies whose orientations are constrained to be in some variety V \u00f0S3\u00dem. Then the map R induces in a natural way a map R\u0302 : V ! \u00f0SO\u00f03\u00de\u00dem. Now suppose that Vj V are some subvarieties. Definition 3.1. V1 and V2 are physically same, if R\u0302\u00f0V1\u00de \u00bc R\u0302\u00f0V2\u00de. In this case we write V1 V2. Moreover if V1 is a proper subvariety of V2 then we write V1-V2. Hence the goal in the analysis is to find the \u2018\u2018smallest\u2019\u2019 or \u2018\u2018most convenient\u2019\u2019 W such that W-V ."
        ]
    },
    {
        "image_filename": "designv11_30_0003292_acc45564.2020.9147392-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003292_acc45564.2020.9147392-Figure2-1.png",
        "caption": "Fig. 2. Trajectories of the vehicles under the control parameters given in Table II.",
        "texts": [
            ", V1 and V5 circumnavigate T1 and T5 with radius 1, etc. In this subsection, we provide a moving-multi-target enclosing example to show that the controller (2) has the potential to capture moving targets although a rigorous analysis is not yet done currently. The targets are initially located at four different positions (4, 4), (4,\u22124), (\u22124,\u22124) and (\u22124, 4), respectively. Using the parameters designed in Table II, eight vehicles are driven to achieve the cooperative multi-target enclosing as shown in Fig. 2(a). After that, from some moment on, all the targets move toward the origin (0, 0) at a speed of 0.04 per second and finally stop at the origin. We can see from Fig. 2(b) that the vehicles keep enclosing and moving along with their targets. When the targets stop at the origin, we can observe that the vehicles soon achieve an even-spaced circular formation around the targets, which is impossible without coordination between each subgroup along the way. 3526 Authorized licensed use limited to: Cornell University Library. Downloaded on September 17,2020 at 10:24:02 UTC from IEEE Xplore. Restrictions apply. EXPLICIT INFORMATION EXCHANGE It is natural in real scenarios when some vehicles in the team loses its observation of their targets, for example, due to obstacle occlusion, limited measurement range, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure3-1.png",
        "caption": "Fig. 3 Snake Manipulator Model",
        "texts": [
            " The elastic intervertebral disc can transform the kinetic energy to elastic potential energy during the movement to reduce energy consumption. Although the motion between adjacent vertebrae is small, the multiple series of vertebrae are accumulated together to form a large range of motion, including flexion and extension, lateral flexion, and rotation motion, which can generate high flexibility to spine organisms. Referring to the biological snake\u2019s spine, the snake manipulator prototype is designed with five segments and 10 DOFs, as shown in Fig. 3. In the figure, 2 DOF motion of each bionic segment is achieved through a universal joint and 3 ropes, where the ropes are applied to imitate the stretching of the biological tendon. The bionic vertebral segment is mainly composed of an upper plate, a lower plate, 3 support columns, and a universal joint. On the upper and lower plates, there are holes for installing ropes. The ropes for driving the vertebral segment are fixed on the lower joint plate, and the ropes for driving the previous segment are guided to the base pose through the mounting holes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003225_aer.2020.54-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003225_aer.2020.54-Figure1-1.png",
        "caption": "Figure 1. Sketch of the stratospheric airship.",
        "texts": [
            " 3) For convenience, with \u03b6 (t) = \u03b2\u22121\u03b2\u0307, it is easy to achieve the expression of \u03b6 ( j)(t)( j = 0, 1, . . . , n \u2212 1) through mathematical methods. Moreover, we can obtain \u03b6 ( j), which belongs to the class Cn\u22121\u2212j function of t and are bounded everywhere. 3.0 PROBLEM FORMULATION 3.1 Stratospheric airship model The main structure of the stratospheric airship investigated herein contains a teardrop-shaped helium balloon, aerodynamic control surfaces, propellers, and an equipment tank, as exhibited in Fig. 1. The equipment tank is suspended from the body, whereas the rudders are attached to the empennage surfaces and provide steering torques. The propellers supply the primary propulsive forces for flight and are fixed on both sides of the equipment tank. Due to the existence of fans and valves, the airship can maintain a stable shape during operation. Thus, we consider the airship as a rigid body and neglect the aeroelastic effects. In this paper, we concentrate on the horizontal motions of the airship and assume that the airship is flying in a buoyancy-weight balance condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003627_rpj-01-2020-0009-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003627_rpj-01-2020-0009-Figure8-1.png",
        "caption": "Figure 8 A component of a surgical instrument produced by additive manufacturing (A, C) and bulk finished using white oxide blasting, centrifugal finishing and glass bead blasting (B, D). The component is a finger operated lever measuring 12 mm x 30 mm.",
        "texts": [
            " Consideration should also be given to surface roughness differences between horizontal, vertical and inclined surfaces of different angles (Figure 2). Surface finishing Alasdair Soja, Jun Li, Seamus Tredinnick and TimWoodfield Rapid Prototyping Journal An example of a surgical instrument component, a lock lever, was used to test a selection of the surface finishing processes alongside the sample coupons. One lock lever underwent white oxide vapour blasting, CF and a final glass bead blast (the proposed finishing process). Sample photos of this part are shown in Figure 8. The inclusion of these components demonstrates how real parts finished by this process would appear. Further investigation should be performed to evaluate the effect of the post-processing on the mechanical properties, fatigue behaviour and corrosive properties of AM instruments. This study systematically investigated a mass surface finishing strategy for additively manufactured parts for end-use surgical instruments. The strategy focused on reducing and optimising the surface roughness of specifically designed stainless steel AM coupons to closely match those of machined surface finishes accepted in current surgical instrument manufacture"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003256_1.c035739-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003256_1.c035739-Figure8-1.png",
        "caption": "Fig. 8 Axis deviation method.",
        "texts": [
            " Meanwhile, in the stress level test, strain rosettes are arranged at the locations of interest to derive the stress levels at the corresponding parts. The challenge lies in the measurement of friction torque of hole\u2013shaft fit. In the present test, the strain rosettes are arranged on the two symmetry planes of 0 and 180 deg angles on the shaft pin. The intermediate gauge is required to coincide with the pin neutral layer, whereas the rest of the gauges form 45 deg angles with the neutral layer, as shown in Figs. 6 and 7. Friction torque is obtained by eliminating the influence of other external loads by bridging means. As shown in Fig. 8, to simulate the axis deviation of the landinggear retractionmechanismcaused bywing deformation and assembly error, the position of up-node rotation axis on the landing gear is adjusted during the test. The specific implementation procedure of axis deviation is as follows: four sets of wedge-shaped plates (as shown in Figs. 9 and 10) are arranged, where each of the large and small plates is matched into one set. The small wedge-shaped plates are weld connected with the up-node fixture, whereas the large wedge-shaped plates are attached to the bevels of small wedgeshaped plates and locked by bolts",
            " To release the radial DOF, the kinematic pair relationship between shaft pin and hole is defined as the point-to-surface constraint. Regarding selection of friction parameters, the semilubricating dynamic friction coefficient between the steel shaft pin and the copper bushing is set as 0.1, and the critical switching velocity between dynamic and static friction is 0.5 mm\u2215s. During simulation, the axis deviation is made to stagger the intersection point by moving the axis of upper node, as shown in Fig. 8. Contact force is simulated with the popular L\u2013N model [16]. The model simplifies the contact impact into a spring damp system, in which the stiffness term is a nonlinear function (classical Hertz force model) with the embedded depth of impact body. Besides, a damping term associated with restitution coefficient is introduced to simulate the energy loss during contact using damping force. Accordingly, the normal contact force is expressed as FN K\u03b43\u22152 1 3 1 \u2212 c2e 4 \u03b4 0 \u03b4 0 \u2212 (1) The contact stiffness coefficient of impact body K is correlated with the material properties and structural shapes of contact planes, which can be expressed as K 4 3\u03c0 hi hj RiRj Ri Rj 1\u22152 (2) The equivalent spherical radii Ri; Rj for the impact between two bodies can be indicated by Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002467_1.4964104-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002467_1.4964104-Figure6-1.png",
        "caption": "Fig. 6. Setup and free-body diagram for the measurement of the maximum frictional force for different normal forces.",
        "texts": [
            " For this study, we did not obtain time resolved force versus time data and we need to make some simplifying assumptions. We ignore any dependence of the contact time on the paddle angle, speed, and type, as well as the speed and spin of the ball. With the rough estimate of contact time, we can estimate normal and frictional forces if we know the coefficient of friction of the rubber sheet. For this, we conducted a classical friction experiment by sliding the ping-pong ball, with different weights added, across the surface of interest (the rubber sheet). Figure 6 shows the experimental setup. The ping-pong ball was placed in a measuring cup and fixed with masking tape so that it could not rotate. The total mass of the tape, ball, and measuring cup was 61 g. A 50-g weight hanger is attached for all but the measurement with the lowest normal force. Additional masses, up to a total of 1,511 g, are added in 100-g increments. The weight hanger is stabilized with minimal vertical force with one hand while the other hand applies a horizontal force that is measured with a force sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001734_978-3-319-10807-0_16-Figure16.3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001734_978-3-319-10807-0_16-Figure16.3-1.png",
        "caption": "Fig. 16.3 A team of Pioneer 3-DX robots",
        "texts": [
            " Figure16.2 shows the floor plan and the extracted topological map on top of the 67.85 \u00d7 26.15m environment. The resulting topology is a noncomplete, connected and sparse graph, like most real world environments. In these tests, robots must overcome noisy sensor readings, localization issues, and even robot failures, which are usually ignored or not precisely modeled in simulation experiments. Therefore, a team of three Pioneer-3DX robots equipped with an Hokuyo URG-04LX-UG01 laser was used, as seen in Fig. 16.3. The Pioneer 3-DX is a lightweight two-wheel differential drive robot for indoor use, equipped with two high-speed, high-torque, reversible-DC motors. Each motor has a high resolution optical quadrature shaft encoder for precise position, speed sensing, and advanced dead-reckoning. These robots are highly popular due to their versatility, reliability, and durability. They can operate continuously for 8\u201310h, with a maximum load of 23kg on top of the platform. In terms of dimensions, the robot has a diameter of 45"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001125_s12206-013-1181-9-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001125_s12206-013-1181-9-Figure3-1.png",
        "caption": "Fig. 3. A schematics of mooring system for dynamic analysis with active cable.",
        "texts": [
            " The numbers of winches installed depends on the vessel capacity. For convenience, a mooring system with a single winch is considered as shown in Fig. 2 because it is possible to extend the result obtained from this study to the general mooring system. Here, we assume that a single winch is installed on the right side (active cable) of the vessel and the cable on the opposite side (passive cable) is just anchored to the seabed without a winch. It means that the vessel is moved on the horizontal plane and controlled by a single winch operation. Fig. 3 depicts a schematic diagram of the controlled system. The parameters shown in Fig. 3 are defined as: sM : Mass of the vessel sK : Spring coefficient sD : Damping coefficient x : Displacement of the vessel from the initial point f : Input force to the vessel (disturbance force is included) ( 1,..., )im i p= : Particle mass of cable ( 1,..., )ik i p= : Spring coefficient of each particle mass ( 1,..., )id i p= : Damping coefficient of each particle mass ( 1,..., )ix i p= : Displacement of each particle mass. The dynamic equations of the controlled system shown in Fig. 3 are described as follows: 1 1 1 1( ) .s s s w dM x D x K x b x k x x f f f+ + + + - = = +&& & & (1) 1 1 1 1 1 1 2 1 2 2 2 2 2 2 2 1 3 2 3 1 1 1 1 1 1 2 1 1 1 ( ) ( ) 0 ( ) ( ) 0 ( ) ( ) 0 ( ) 0. p p p p p p p p p p p p p p p p p p p m x d x k x x k x x m x d x k x x k x x m x d x k x x k x x m x d x k x x k x - - - - - - - - - + + + - + - = + + - + - = + + - + - = + + - + = && & && & M && & && & (2) where Eq. (1) describes the vessel dynamic equation and Eq. (2) is the active cable dynamics. However, in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003503_j.ijpvp.2020.104221-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003503_j.ijpvp.2020.104221-Figure5-1.png",
        "caption": "Fig. 5. Grid partition graph of tool joint.",
        "texts": [
            " International Journal of Pressure Vessels and Piping 188 (2020) 104221 The structure drawing and design parameters are shown in Fig. 3 and Table 1. For ensuring calculation accuracy and speed at the same time, reduced integration Hexahedral elements C3D8RT are chosen in this model. Grid sensitivity analysis is carried out. The result can be seen in Fig. 4. When the number of elements is larger than 398520, the ultimate working torque changes little. Thus, the model with 398520 elements is chosen in this paper (Fig. 5). Compared with the results in Chen\u2019s article [25], the accuracy of the present model is given. The ultimate working torque is 44.85 kN m by using finite element model of this paper (Fig. 6). The error doesn\u2019t exceed 4%. (the ultimate working torque in Chen\u2019s article [25] is 43.213 kN m). 35HM steel is selected in the analysis [30]. Specific heat and conductivity change little with temperature and have little effect on the mechanical behavior of tool joint. So these parameters are assumed to be constant, which are shown in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003914_gt2015-43940-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003914_gt2015-43940-Figure7-1.png",
        "caption": "Figure 7. CAD image of rig used to measure pressure drop and heat transfer in additively manufactured test coupons.",
        "texts": [
            " Despite the variation in roughness between different surfaces in the same channel, the area weighted average roughness of all coupons was similar in range (9-14 \u03bcm) except for the L-2x-Co coupon. The reason for the difference may be due to a combination of factors such as the material, the machine parameters and the geometry of the coupon. Even with this variation, much of the roughness data presented here is similar in magnitude to the results many have found regarding effects of orientation and overall roughness magnitude [2\u20134,6]. A test rig was built to collect pressure drop and heat transfer measurements of each of the test coupons. The rig (shown in Figure 7) was built with a smooth contraction chamber that supplied air to the coupon inlet. The inlet contracted the crosssectional flow area by a factor of 11 for the coupon with the most flow area which ensured a uniform velocity just before the air entered the coupon. The exit expansion chamber was made identical to the inlet for simplicity of fabrication. Pressure taps were installed upstream of the inlet contraction and downstream of the exit expansion to measure the pressure drop across the coupon"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001380_icra.2014.6907727-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001380_icra.2014.6907727-Figure5-1.png",
        "caption": "Fig. 5. The schematic of the experiment that was performed for out-ofplane detection of the cannula tip",
        "texts": [
            " To determine the image quality and spatial resolution of the transducer, the object shown in Fig.4a was scanned along its width. The dimensions of the cross-section is given in Fig. 4b. Fig. 4d shows several dimensions of the object measured using the cursor in the ultrasound console. The maximum error in dimensions was measured to be 0.015 cm. This implies that the spatial accuracy of in-plane tracking is limited by the spatial accuracy of the ultrasound. 0 .6 The ultrasound transducer is placed at right angles to the cannula as demonstrated in Fig. 5. The imaging depth was set to 8 cm. The initial position of the transducer is away from the cannula tip such that the cannula cannot be seen in ultrasound images. The transducer is moved towards the cannula tip with 0.127 mm increments and a total number of 185 ultrasound images were recorded. To detect the tip of the cannula, Hough circle transform was implemented. The ultrasound images were pre-processed using thresholding. Downscaling, upscaling, dilation, and erosion operations followed by blurring were applied recursively to obtain a clear image that is free of any noise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000500_12.2186385-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000500_12.2186385-Figure1-1.png",
        "caption": "Figure 1. Schematic representation of the printing process consisting of screen-printing of a silver ink layer for interconnection a), a PEDOT:PSS which acts as channel b), a layer of silver/silver chloride, PEDOT:PSS or polyaniline as gate material c) and a dielectric layer d).",
        "texts": [
            " The use of printing techniques to obtain OECTs implies some advantages compared to standard pholithographic process. It allows indeed a faster manufacturing with low cost techniques. Thus, getting dedicated geometries and materials to a specific application is straightforward. PEN TEONEX Q65FA from Dupont Teijin was used as a substrate. A first layer of silver paste Metalon HPS-021LV from NovaCentrix was deposited by screen printing for electrical connexions, was annealed at 100\u00b0C for 30min and had a thickness around 6\u00b5m (Fig 1.a). Then a first layer of PEDOT:PSS Clevios SV4 from Heraeus was deposited by screenprinting and annealed at 130\u00b0C for 30min (Fig 1.b). The final thickness of the PEDOT:PSS layer is around 500nm. Following this, the gate material is deposited by screen printing, depending on the device (Fig 1.c). A layer of Ag/AgCl 5064 (Dupont) was deposited for high transconductance OECTs, and annealed at 120\u00b0C for 4min. A PEDOT:PSS layer with the same protocol as previously described was deposited for biosensing applications and a polyaniline (PANIPLAST from Rescoll) layer for pH sensor application was screen-printed, annealed at 100\u00b0C for 7min. The final layer was a screen-printed dielectric based poly(vinylidene fluoride) (PVDF) cured at 100\u00b0C for 5min (Fig 1.d). Proc. of SPIE Vol. 9568 95681E-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 10/20/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx The shape of the different electrodes allows to select the better transconductance at a given gate potential. It has been shown by Rivnay et al.[16], that increase significantly the transconductance could be obtained by changing the width-tolength ratio and thickness, as a result we will choose here a four to one gate to channel ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003251_j.matpr.2020.06.456-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003251_j.matpr.2020.06.456-Figure1-1.png",
        "caption": "Fig. 1. The chip encapsulation during",
        "texts": [
            " In this paper, a newmethodology of encapsulating 550-mmRFID sensor (Murata LXMSJZNCMF-198) [32] via a screw extrusion process is proposed. The fabricated filament with a diameter of 3 mm is produced using a 1.5 mm nozzle specifically designed during this project and controlled by 200mW UHF RFID reader. The final product is then printed by an FDM 3D printer. To realize the new technique, we made some modifications to both the filament production and the FDM 3D printing process. The extrusion machine was modified (see Fig. 1) and equipped with a unique die unit and pneumatic gripper designed for new microchipped fil- Please cite this article as: M. Pekgor, M. Nikzad, R. Arablouei et al., Sensor-ba Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.06.456 ament manufacturing process (see Fig. 2). For the FDM 3D printing machine, also an original nozzle was designed and fabricated. A set of new steps were introduced to the 3D printing process controlled by hardware and software. The technical specifications of the encapsulated RFID sensor are shown in Table 1",
            " Table 1 Technical Specifications of Murata LXMSJZNCMF-198 including Impinj Monza R6 Tag chip [33]. Extended operating temperature 40 C to 85 C Storage temperature 40 C to 85 C Assembly survival temperature 150 C (for one minute) Temperature rate of change 4 C /sec Electrostatic discharge 2000 V (Human body model) related to the junction temperature of the chip, is crucial for most of silicon technologies [35]. In the new encapsulation technique, RFID-encapsulated filaments were produced by using a pneumatic gripper injection unit at the end of twin-screw extrusion unit (see Fig. 1 and Fig. 2). The used pneumatic gripper is made of brass and includes a pneumatic piston and a needle control by a programmable logic control (PLC) device. It can also be mounted on the robotic arms. While being poxy cyclo hexyl methyl3,4epoxycyclohexanecarboxylate) (a) molding, (b) the final ould one by one separation process by drilling, (b) the final product. Fig. 5. RFID Microchips coated with epoxy resin and blended with polymer (PP) sintered at 180 C twenty minutes in the vacuum oven (a) moulding, (b) vacuum oven"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001740_s1560354715040024-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001740_s1560354715040024-Figure2-1.png",
        "caption": "Fig. 2. Decomposition of the function fA into the basis e1, e2.",
        "texts": [
            " The absolute value of the vector \u03b1A is determined by the forces applied to the slider and is not greater than one. This approach was previously proposed by John H. Jellett [11, 12]. Since the point O is fixed, the vectors \u03c5A are directed perpendicularly to the radius vector rOA. Suppose that the vectors \u03b1A have the same orientation. Assume also that the absolute value of the vector fA REGULAR AND CHAOTIC DYNAMICS Vol. 20 No. 4 2015 is invariant with respect to the point A: fA \u2261 f . Combining the vectors \u03c5\u22121 A \u03c5A and \u03b1A in one formula, we obtain the following expression (Fig. 2): fA = f\u03bee1 + f\u03b7e2, f\u03be = \u2212f \u03b7 \u221a \u03be2 + \u03b72 , f\u03b7 = f \u03be \u221a \u03be2 + \u03b72 , \u22121 f 1. The continuity condition of the contact between the body and the surface imposes three independent constraints on the kinematic characteristics. Therefore, the model of normal stresses nA should include three independent parameters \u03bb0, \u03bb1 and \u03bb2, which are determined from these limitations each time. This model of normal stress distribution is dynamically consistent [11] nA = \u03bb0 + \u03bb1\u03be + \u03bb2\u03b7, (\u03be, \u03b7) \u2208 S. (3.6) The physical interpretation of this formula can serve as a representation of small deformations of the surface in the contact area, leading to the normal stresses by Hooke\u2019s law"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003699_icem49940.2020.9270705-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003699_icem49940.2020.9270705-Figure1-1.png",
        "caption": "Fig. 1. Six-phase 12/10 SRM prototype. (a) Wound stator. (b) Rotor.",
        "texts": [
            " Afterwards, a Direct Energy Control (DEC) strategy is proposed based on the DTC A Direct Energy Control Technique for Torque Ripple and DC-link Voltage Ripple Reduction in Switched Reluctance Drive Systems Xu Deng and Barrie Mecrow S Authorized licensed use limited to: University of Ghana. Downloaded on December 19,2020 at 10:28:46 UTC from IEEE Xplore. Restrictions apply. strategy. Compared to the conventional control strategy, the DC-link voltage ripple and the torque ripple are both reduced significantly with the proposed strategy. It is well understood that increasing the number of phases reduces torque ripple. Therefore, a six-phase 12/10 SRM is proposed in Fig. 1. Table I gives the design parameters for this machine. Asymmetric Half Bridge (AHB) converters are the most popular choice for SRM drives as they give independent control of each phase. Fig. 2 shows a six-phase AHB power inverter which is employed for this six-phase SRM drive system. A dynamic model of a six-phase SRM is developed in the MATLAB/SIMULINK environment for simulation. The model used electromagnetic parameters from the prototype six-phase, 12/10 SRM, and the parameters are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002645_3029610.3029619-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002645_3029610.3029619-Figure3-1.png",
        "caption": "Figure 3. The Kinematics scheme for inverse problem [20].",
        "texts": [
            " The Raspberry PI reads the accelerometer data via I2C interface, interprets it and sends commands to quad-copter via WIFI according to hand movement. On the top of the Raspberry PI case it is mounted a board with two buttons, one for take-off and one for landing. The quadcopters are very robust, because they have only four brushless motors attach to a propeller. In order to fly, the mechanical simplicity, needs to be compensated with advance controls algorithms. In the literature are presented studies quite advanced for the UAVs (Unmanned aerial vehicles), and quadcopters. In Fig. 3 it is noted with \u03be the inertial frame x, y, z axis which define the absolute linear position of the quadcopter. With \u03b7 it is marked the angular position defined by the three angles: \u03b8 known as pitch angle (rotation of the quadcopter around y axis), \u03c6 known as roll angle (rotation of the quadcopter around x axis), \u03c8 known as yaw angle (rotation of the quadcopter around z axis) The equations (1) define the linear and angular vectors.            Z Y X                          q (1) Equations (2) show the linear velocity VB and angular velocity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002236_j.ifacol.2016.07.146-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002236_j.ifacol.2016.07.146-Figure1-1.png",
        "caption": "Fig. 1. 2D Plotter, (i) main construction, (ii) x-axis linear table, (iii) y-axis rail, (iv) plotting head, (v) microcontrollers Arduino Uno and Nano, (vi) controllers for stepper motors.",
        "texts": [
            " The design of the construction was determined by the main goals, i.e., user-friendly handling and budget limitations. Therefore we tried to minimize the overall size and weight of the 2D Plotter to simplify its utilization. This allows the device to become a portable piece of educational equipment that can be easily transferred between laboratories. The portability of the device also brings possibilities to promote the control engineering education in secondary schools and other institutions that prepare students for practice and higher education. Figure 1 shows the overall physical construction of 2D Plotter. The device has following structural, operational and electronic parts: \u2022 Main frame \u2013 The main construction that holds the device together and allows the attachment of sensors, actuators and electronics is build on the OpenBeam platform 1 , an extruded aluminum framing system brought to the public by a successful campaign at Kickstarter. The OpenBeam provides a modular aluminum frames of different lengths and various types of brackets. This allows to build almost any kind of mechanical construction",
            " First is the Atmel ATMega328p, the 8-bit AVR microcontroller which is a part of development board Arduino UNO Rev3 2 and second one is the ATMega168 on Arduino Nano 3. The essential task was to ensure accurate operating of the device parts in full 3 dimensional space in order to control a pen for the plotting task, or a light sensor for the optimization task. This complex task was expanded into three particular tasks provided by two stepper motors for movements in x-axis and y-axis and one servomotor for vertical movement of main 2D Plotter \u2019s tool. The movement in x-axis is ensured by linear table attached on rails (Fig. 1, (ii)). This construction was chosen to simplify the development of device as well as to achieve the high operational precision. Many similar devices use the attachment of linear actuators on each other, however such approach has its own limitations especially in the cable organization and mechanical clearance of head control. The size of the table is 120\u00d7 150mm and it moves on the rails with a diameter 6mm. The exact positioning of the table is ensured by stepper motor to which the table 2 http://arduino.cc/en/Main/arduinoBoardUno 3 http://arduino.cc/en/Main/arduinoBoardNano is attached via a belt. The table serves as the surface for the placement of a paper or tablet. The movements in y-axis are managed in the similar manner by means of rail perpendicular to the axis of table movement (Fig. 1, (iii)). This movement is also ensured by a belt driven by stepper motor. This rail is a guide for the main head that carries the plotting tool or light sensor. Finally, the movements in z-axis are quite limited compared to the movements in x- and y-axis. The z-axis is managed considering just for the binary values, that stand for active or inactive position of the main tool (Fig. 1, (iv)). The z-axis is handled by servomotor that moves the main tool down to the surface and then it is returned back into the original position by a spring. The 2D Plotter uses several electronic devices in order to ensure proper operation. As mentioned earlier, the main control component of the device is microcontroller board Arduino Uno (Fig. 2). The board contains an electric signal interface that provides 14 digital I/O pins, six of which can be used as Pulse-Width Modulation (PWM) outputs, additional six analog input pins with 10-bit resolution, and communication interfaces such as Universal Asynchronous Receiver/Transmitter (UART), Serial Peripheral Interface (SPI) and Inter-Integrated Circuit (I2C)",
            " Moreover, the price of this microcontroller is quite acceptable. Control unit is able to provide just a limited power that is not sufficient to ensure the movements of all attached motors. Therefore, it was necessary to implement two stepper motor drivers TB6560 for the movements of the table and head in x- and y-axis, respectively. Particularly, TB6560 (Fig. 3) accepts the PWM signal 2016 IFAC ACE June 1-3, 2016. Bratislava, Slovakia Juraj Oravec et al. / IFAC-PapersOnLine 49-6 (2016) 016\u2013021 19 Fig. 1. 2D Plotter, (i) main construction, (ii) x-axis linear table, (iii) y-axis rail, (iv) plotting head, (v) microcontrollers Arduino Uno and Nano, (vi) controllers for stepper motors. controllers which allow to perform the micro-steps up to 1/16 of step size. For the vertical movement of the main tool (drawing tool or light sensor) the servomotor is used. The main task of this motor is either to push the drawing tool against the surface or to change the vertical position of light sensor over the tablet in order to achieve the best reading performance",
            " First is the Atmel ATMega328p, the 8-bit AVR microcontroller which is a part of development board Arduino UNO Rev3 2 and second one is the ATMega168 on Arduino Nano 3. The essential task was to ensure accurate operating of the device parts in full 3 dimensional space in order to control a pen for the plotting task, or a light sensor for the optimization task. This complex task was expanded into three particular tasks provided by two stepper motors for movements in x-axis and y-axis and one servomotor for vertical movement of main 2D Plotter \u2019s tool. The movement in x-axis is ensured by linear table attached on rails (Fig. 1, (ii)). This construction was chosen to simplify the development of device as well as to achieve the high operational precision. Many similar devices use the attachment of linear actuators on each other, however such approach has its own limitations especially in the cable organization and mechanical clearance of head control. The size of the table is 120\u00d7 150mm and it moves on the rails with a diameter 6mm. The exact positioning of the table is ensured by stepper motor to which the table 2 http://arduino.cc/en/Main/arduinoBoardUno 3 http://arduino.cc/en/Main/arduinoBoardNano is attached via a belt. The table serves as the surface for the placement of a paper or tablet. The movements in y-axis are managed in the similar manner by means of rail perpendicular to the axis of table movement (Fig. 1, (iii)). This movement is also ensured by a belt driven by stepper motor. This rail is a guide for the main head that carries the plotting tool or light sensor. Finally, the movements in z-axis are quite limited compared to the movements in x- and y-axis. The z-axis is managed considering just for the binary values, that stand for active or inactive position of the main tool (Fig. 1, (iv)). The z-axis is handled by servomotor that moves the main tool down to the surface and then it is returned back into the original position by a spring. The 2D Plotter uses several electronic devices in order to ensure proper operation. As mentioned earlier, the main control component of the device is microcontroller board Arduino Uno (Fig. 2). Fig. 2. Microcontroller board Arduino UNO equipped with the 8-bit AVR microprocessor ATMega328p. The board contains an electric signal interface that provides 14 digital I/O pins, six of which can be used as Pulse-Width Modulation (PWM) outputs, additional six analog input pins with 10-bit resolution, and communication interfaces such as Universal Asynchronous Receiver/Transmitter (UART), Serial Peripheral Interface (SPI) and Inter-Integrated Circuit (I2C)",
            " Various graphical editors allow to export vector graphics to G-code (e.g. Inkscape 6 with G-code extension). In this way, students can draw their own vector objects in graphical editor and export the resulted G-code to be used by Grbl Controller. Fig. 5 shows the GUI of Grbl Controller with imported G-code representation of drawing. User can see both, the generated curves for plotting tool as well as the list of instructions for particular movements. The final result of plotting task can be seen in Fig. 1. The second possibility for using and controlling the 2D Plotter at IAM is use of MATLAB as the operational 5 http://zapmaker.org/projects/grbl-controller-3-0 6 https://inkscape.org 2016 IFAC ACE June 1-3, 2016. Bratislava, Slovakia 19 20 Juraj Oravec et al. / IFAC-PapersOnLine 49-6 (2016) 016\u2013021 software, since MATLAB is widely used in the education process at our institute. To extend the usage potential of device, an MATLAB based Application Programming Interface (API) has been developed. This API provides functionality for direct control of device\u2019s actuators and allows to collect data from attached sensors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002827_j.optlaseng.2020.106065-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002827_j.optlaseng.2020.106065-Figure12-1.png",
        "caption": "Fig. 12. The state of the particle and the surrounding fluids at 6 \ud835\udf07s: (a) where the powder particle is located; (b) speed distribution of the particle and the surrounding fluids; (c) the speed distribution of the pool around the particle.",
        "texts": [
            " At the same time, due to the increase n the contact area between the particle and the Ti-6Al-4V fluid, the surace tension continues to increase, which causes the powder particle to apidly reduce its running speed. With the decrease of the running speed f the particle to ~1.6 m s \u2212 1 at 29 \ud835\udf07s, the reaction force of the fluid on he particle is reduced, and the contact area between the particle and he Ti-6Al-4V fluid increases much slower due to the generation of the ap between the upper part of the particle and the Ti-6Al-4V fluid (as hown in Fig. 10 at ~29 \ud835\udf07s), and thus leads to a slower decrease of the unning speed of the particle. Fig. 12 shows the process of the powder particle entering the melt ool at 6 \ud835\udf07s, where Fig. 12 (a) is the location of the powder partile. Fig. 12 (b) shows the longitudinal section speed distribution of the owder particle taken along the center of the particle, which reveals c i c p ~ t t t s ~ a i o w s F c t t s a i t f f s p g 2 t p l w o t w z p p u n p v c c T i t i ~ learly the movement speed of the powder particle and the surroundng fluids, and Fig. 12 (c) is the speed distribution of the melt pool. As an be seen from the figure, when the particle starts to enter the melt ool, the speed of the particle did not change substantially, which is 2.25 m s \u2212 1 . The gas-liquid interface remains basically horizontal, and he surface of the melt pool is not significantly disturbed, as the flucuation area on the gas-liquid interface is less than twice the cross secion of the particle. However, the melt pool is subject to a relative high peed and speed gradient around the particle, as the peak speed reaches 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002349_978-3-319-45781-9_26-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002349_978-3-319-45781-9_26-Figure2-1.png",
        "caption": "Fig. 2. Pliers structure.",
        "texts": [
            " A kinematical chain has been defined as follows: 3 rotating pairs with parallel lines of actions are used for the mechanical arm movements; 2 rotating pairs with perpendicular lines of action are used for the pliers movements and 2 gearwheels are used for pliers opening and closing. The main selection parameters are precision and ease positioning, component resistance, reduced amount of necessary space, easy architecture. The preliminary design of the mechanical arm and of the plier has been developed, introducing the selected kinematical chains (Figures 1 and Figure 2). The aim of the research activity is to optimize the sizing of the rotating pairs to be manufactured as unibody structures, in order to allow the movements within the joint (clearance effect), with a proper positioning between all the parts and with the feasibility to make all the possible positioning of the arm extremity as stable equilibrium position (posable effect). The final solution is prototyped with a FDM \u2013 Fused Deposition Modeling \u2013 technique as a demonstration of the developed concept"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000967_icarcv.2014.7064591-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000967_icarcv.2014.7064591-Figure1-1.png",
        "caption": "Fig. 1: Schematic description of an induction motor with a single phase stator winding short circuit fault",
        "texts": [
            " Here, the motor life is defined as the stator winding insulation life suffered from thermal stresses because from a thermal point of view, the stator winding insulation is the weakest part of induction motors. A particle filter (PF) method is used to realize joint state and parameter estimation, where the motor model is augmented to include the unknown faulty parameters. Based on the estimated unknown faulty parameters, the RUL of the faulty component can be computed if the failure threshold is defined. Simulation results are provided to validate the proposed method. II. INDUCTION MOTOR MODEL WITH STATOR WINDING SHORT CIRCUIT FAULT Fig. 1 depicts an induction motor with a single phase stator winding short circuit fault. The induction motor model under fault condition in abc frame can be represented as Vs = RsIs + p\u03bbs 0 = RrIr + p\u03bbr \u03bbs = LssIs + LsrIr \u03bbr = LrsIs + LrrIr (1) where p denotes the operator d/dt, stator voltage vector Vs = [vas1 vas2 vbs vcs] T , stator current vector Is = [ias (ias \u2212 if ) ibs ics] T , rotor current vector Ir = [iar ibr icr] T , stator resistor vector Rs = Rsdiag[1\u2212 \u03bc \u03bc 1 1] T , rotor resistor vector Rr = Rrdiag[1 1 1] T , stator flux vector \u03bbs = [\u03bbas1 \u03bbas2 \u03bbbs \u03bbcs] T , rotor flux vector \u03bbr = [\u03bbar \u03bbbr \u03bbcr] T , stator self-inductance matrix Lss = Llsdiag[1\u2212 \u03bc \u03bc 1 1] +Lms \u23a1 \u23a3 (1\u2212\u03bc)2 \u03bc(1\u2212\u03bc) \u2212(1\u2212\u03bc)/2 \u2212(1\u2212\u03bc)/2 \u03bc(1\u2212\u03bc) \u03bc2 \u2212\u03bc/2 \u2212\u03bc/2 \u2212(1\u2212\u03bc)/2 \u2212\u03bc/2 1 \u22121/2 \u2212(1\u2212\u03bc)/2 \u2212\u03bc/2 \u22121/2 1 \u23a4 \u23a6, rotor self-inductance matrix Lrr = [ Llr+Lms \u2212Lms/2 \u2212Lms/2 \u2212Lms/2 Llr+Lms \u2212Lms/2 \u2212Lms/2 \u2212Lms/2 Llr+Lms ] , mutual inductance matrix Lsr = LT sr = Lms [ (1\u2212\u03bc) cos(\u03b8r) (1\u2212\u03bc) cos(\u03b8r+2\u03c0/3) (1\u2212\u03bc) cos(\u03b8r\u22122\u03c0/3) \u03bc cos(\u03b8r) \u03bc cos(\u03b8r+2\u03c0/3) \u03bc cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r) cos(\u03b8r+2\u03c0/3) cos(\u03b8r+2\u03c0/3) cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r) ] , The machine model in (1) can be reformulated as V\u2032 s = RsI\u2032s + p\u03bb\u2032 s + \u03bcA1if 0 = RrIr + p\u03bbr \u03bb\u2032 s = L\u2032 ssI\u2032s + L\u2032 srIr + \u03bcA2if \u03bbr = L\u2032 rsIs + LrrIr + \u03bcA3if (2) where V\u2032 s = [vas vbs vcs] T , I\u2032s = [ias ibs ics] T , \u03bb\u2032 s = [\u03bbas1 + \u03bbas2 \u03bbbs \u03bbcs] T = [\u03bbas \u03bbbs \u03bbcs] T , A1 = \u2212[Rs 0 0] T , A2 = [\u2212(Lls + Lms) Lms/2 Lms/2] T , A3 = \u2212Lms[cos(\u03b8r) cos(\u03b8r + 2\u03c0/3) cos(\u03b8r \u2212 2\u03c0/3)] T , L\u2032 ss = [ Lls+Lms \u2212Lms/2 \u2212Lms/2 \u2212Lms/2 Lls+Lms \u2212Lms/2 \u2212Lms/2 \u2212Lms/2 Lls+Lms ] , L\u2032 sr = Lms [ cos(\u03b8r) cos(\u03b8r+2\u03c0/3) cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r) cos(\u03b8r+2\u03c0/3) cos(\u03b8r+2\u03c0/3) cos(\u03b8r\u22122\u03c0/3) cos(\u03b8r) ] , The voltage and flux linkage equation for the shorted turns as2 can be represented as vas2 = \u03bcRs(ias \u2212 if ) + p\u03bbas2 = Rf if \u03bbas2 = \u2212\u03bcAT 2 I\u2032s \u2212 \u03bcAT 3 Ir \u2212 \u03bc(Lls + \u03bcLms)if (3) In abc frame, the electromagnetic torque can be represented as Te = P 2 I\u2032Ts \u2202L\u2032 sr \u2202\u03b8r Ir \u2212 P 2 \u03bcLmsif [ 3 2 iar sin \u03b8r + \u221a 3 2 (ibr \u2212 icr) cos \u03b8r] (4) In stationary reference frame using dq variables, the ma- chine model in (2) can be expressed as vqs = Rsiqs + p\u03bbqs \u2212 2 3 \u03bcRsif vds = Rsids + p\u03bbds 0 = Rriqr + p\u03bbqr \u2212 \u03c9r\u03bbdr 0 = Rridr + p\u03bbdr + \u03c9r\u03bbqr \u03bbqs = Lsiqs + Lmiqr \u2212 2 3 \u03bcLsif \u03bbds = Lsids + Lmidr \u03bbqr = Lmiqs + Lsiqr \u2212 2 3 \u03bcLmif \u03bbdr = Lmids + Lridr (5) where \u03c9r = p\u03b8r, Lm = 3/2Lms, Ls = Lls + Lm, Lr = Llr + Lm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure3-1.png",
        "caption": "Fig. 3 Modeling of driving with cylindrical driving pin",
        "texts": [
            " M\u00e1ndy Department of Technical Preparatory and Production Engineering, College of Ny\u00edregyh\u00e1za, S\u00f3st\u00f3i u. 9-11, 4400 Ny\u00edregyh\u00e1za, Hungary e-mail: bodzassandor@nyf.hu Power dissipation in the gear can be reduced significantly with these modern drive pairs which are characterized by favorable hydrodynamic conditions, great strength, and high efficiency [4]. In case of power loss, it is important to apply those geometrical characteristics of the cog which result in good connection terms. During our examination, we started from the modeling of the driving of cylindrical driving pin (Fig. 3). We supposed that there is a punctual connection between the driving pin and the lathe fork. That is how we could determine the path of the contact point between the lathe fork and the driving pin during one turning. Figure 4 shows that in the beginning of turning, on the axial intercept plane of the spindle the contact point between the driving pin [I.] and the lathe fork [III.] is point A\u2032. The lathe fork [III.] is situated perpendicular to the worm axis [IV.], namely it encloses half cone angle \u03b41 with the head surface of the spindle [II",
            " 7): rhP \u00bc rPe rm \u00f023\u00de The polar equation of the ellipse section of the driving pin is (Fig. 7): xcs \u00bc rcs sin 8 1 ycs \u00bc rcsen cos 8 1 \u00f024\u00de Using the given equations above, we have prepared our own computer program (Figs. 8, 9, 10, 11, and 12). As the program knows the following pieces of information: the geometrical data of the worm, the driving pin diameter, the distance between the lathe fork and the worm shaft neck, the worm half cone angle, and the distance between the center line of the driving pin and the rotational axis of spindle (Fig. 3), then it calculates and represents the angular veloc- ity fluctuation \u03c9p as a function of spindle angular rotation \u03c6p (Fig. 8). It takes into consideration the width of the tolerance level of the pitch and then it also represents pitch error fluctuation as a function of spindle angular rotation \u03c6p (Fig. 9). The program also shows the path curve of the driving pin and the difference between the path of the driving pin as compared to a circle path (Fig. 10). Moreover, it can also represent the contact points of the driving pin and the lathe fork (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001654_acc.2013.6579901-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001654_acc.2013.6579901-Figure1-1.png",
        "caption": "Fig. 1. Top view of the robotic fish undergoing planar motion.",
        "texts": [
            "00 \u00a92013 AACC 591 In Section III we review the classical averaging method and the geometric first-order averaging method, and discuss their pros and cons. The proposed control-oriented, data-driven averaging method is presented and validated in Section IV. Finally, concluding remarks are provided in Section V. In this section we provide a summary of the dynamic model for a tail-actuated robotic fish that is presented in [20]. A diagram of the system, consisting of a rigid body and an active rigid tail, is shown in Fig. 1. Both the body and the tail are assumed to be neutrally buoyant, and we are interested in the planar motion of the system. [X ,Y,Z] denote the inertial coordinates, and [x,y,z] denote the body-fixed coordinates. The position and orientation of the robot\u2019s center of mass relative to the initial coordinates are denoted by [X ,Y,\u03c8 ]. The velocity and angular velocity of the robot expressed in the body-fixed coordinates are denoted by V = [u,v]T and \u03c9 , respectively. Here u and v denote the surge and sway velocities, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure16-1.png",
        "caption": "Figure 16. Velocity field contours for drum diameter 550 mm.",
        "texts": [],
        "surrounding_texts": [
            "Table 1 and the graph (Figure 14) show the calculations of the maximum speed of the velocity fields at fixed drum and beater revolutions 640 rpm and 2100 rpm, respectively, and different values of the drum diameter. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 At drum diameters less than 550 mm, the maximum speed values are preserved, since the maximum flow rates in these cases are already set by the beater and remain unchanged. Figure 14. Velocity field contours for drum diameter 350 mm. Figure 15. Velocity field contours for a drum diameter of 450 mm. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0001303_1.4024087-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001303_1.4024087-Figure2-1.png",
        "caption": "Fig. 2 Finite element model of the NASA GRC gearbox: (right) part of the top cover and front and side walls removed to reveal the inner shafting and gear blanks, and (bottom) enlarged view of the shafting and gear blanks",
        "texts": [
            "-lb), which corresponds to the theoretical minimum of the GTE for this gearset). The simulated levels are normalized to a unit GTE so that actual operating levels may be computed given a known GTE amplitude. The differences between the simulated levels with traditional REBs and the JBs are compared to the differences observed in the NASA GRC measurements. 3.1 Finite Element Modeling of Gearbox Housing. The housing, base structure, and shafts were modeled with quadratic, tetrahedral, and hexahedral finite elements, as shown in Fig. 2. The model includes about 190,000 nodes and 93,000 solid elements. The feet of the base structure were rigidly attached to the ground. The top plate is rigidly connected to the housing at Table 3 Journal bearing parameters Fluid viscosity l\u00bc 0.0294 Pa s Fluid density q\u00bc 980 kg/m3 Journal diameter D\u00bc 0.03195 m Radial clearance C\u00bc 15 lm Bearing length L\u00bc 0.019882 m Rotation speed X\u00bc 50 Hz Journal of Vibration and Acoustics JUNE 2013, Vol. 135 / 031012-5 Downloaded From: http://vibrationacoustics.asmedigitalcollection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000294_rcar47638.2019.9044123-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000294_rcar47638.2019.9044123-Figure1-1.png",
        "caption": "Fig. 1. A robot is manipulating a DLO. The shape of the DLO is represented as a series of feature points, and the positions of feature points are measured with a camera.",
        "texts": [
            " Second, compared with the existing works on the task sequencing [19], [20], the control objective is specified as a series of dynamic regions in this paper, which can effectively addresses the conflicting movement as other points are allowed to move freely inside the corresponding regions to suit the movement of the point being manipulated. The stability of the closed-loop system is rigourously proved with Lyapunov methods, and experimental results are presented to show the performance of the proposed controller. Consider a robotic manipulation system shown in Fig. 1. The robot grasps one end of the DLO (where the other end is fixed) and deforms it into a desired shape by controlling the movement of end effector, and the positions of multiple feature points along the DLO are measured with sensors (e.g. camera) and employed in the robot controller. 978-1-7281-3726-1/19/$31.00 \u00a9 2018 IEEE 840 In general, the velocity of a specific feature point can be related with the velocity of the robot end effector as x\u0307i = Ji(xi, r)r\u0307, (1) where xi=[xi1 , xi2 , \u00b7 \u00b7 \u00b7 , xik ] T\u2208 k is the position of the ith feature in sensory space (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure10-1.png",
        "caption": "Figure 10. Velocity contours for modified body geometry.",
        "texts": [],
        "surrounding_texts": [
            "An optimization calculation was carried out to assess the effect of the hull geometry. The original body geometry (see Figure 1) has been changed by moving the top down 70 mm - Figure 9. Figures 10 - 12 show the contours of the velocity fields and the velocity vectors for the modified housing shown in Figure 9. The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 31.2 m/s. For this geometry, a more even distribution of the air velocity fields is observed. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure13-1.png",
        "caption": "Fig. 13. Beveloid gear pair.",
        "texts": [
            " Presented in Table 2 is evolutoid data for the axial position (along the common generator) w = 89 . 8525 mm. This example considers a beveloid gear set, a summary of the beveloid gear data is presented in Table 3 . This data set is based on a an existing beveloid data set [22] . This example closely parallels a classical helical gear set, yet is enough different to help illustrate the transition from the general spatial involutoid for hypoid gears to the classical planar involute universally used in spur and helical gear sets today. Presented in Fig. 13 is a virtual model of the beveloid gear pair. Presented in Fig. 14 are two beveloid teeth along with an evolutoid and its involutoid. The first row of images are the input tooth for CW and CCW pinion rotation, the second row of images are the input tooth with the involutoid hyperboloid surface for CW and CCW pinion rotation, and the third row of images are the input and output gear teeth in mesh for CW and CCW pinion rotation. The line of action intersects the transverse curves and the two evolutoids (base circles)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure13-1.png",
        "caption": "Fig. 13 Driving Process in Simulation",
        "texts": [
            " The control framework sends information including the driving velocity and simulation time to the virtual prototype, while receives data of the rope length and rope velocity from the virtual prototype. The virtual prototype consists of the geometric and dynamic model of the robot, and can provide the data of rope length, rope velocity, joint angle, joint velocity and endeffector pose in simulation. Due to using motion drive mode, multi-body system simulation will not be possible under redundant drive conditions. Therefore, the ropes are connected with the virtual springs, as shown in Fig. 13. The virtual spring can ensure that the simulation is carried out normally with the reduandant speed driving, and at the same time, the imitate the interferes caused by the mecnanical flexibility. The detailed parameters of experimental prototype in Fig. 14 are shown in Table 2. The arm is controlled through the screw pulling the rope, and the variation of rope length is equal to the linear movement of the ball-screw pair. In order to feedback the posture of the mechanism, a photoelectric encoder is generally installed at the joint for measuring the joint angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002850_s00170-020-05095-2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002850_s00170-020-05095-2-Figure3-1.png",
        "caption": "Fig. 3 Finite element model",
        "texts": [
            " Take the interface between middle bearing\u2019s outer ring and its bearing sleeve as an example; the constraints of variables in formula 16 can be determined as follows. s:t: 100\u2264H \u2264350 40\u2264k1\u226450 40\u2264k2\u226450 0:8\u2264\u03c31\u22641 0:8\u2264\u03c32\u22641 0:02\u2264\u03b5\u22640:04 0:25\u2264\u03bc1\u22640:3 0:25\u2264\u03bc2\u22640:3 175\u2264E1\u2264220 175\u2264E2\u2264220 8>>>>>>>>< >>>>>>>>: \u00f017\u00de The radius of the bearing and its sleeve is determined by the structure of the spindle. Using the constrained optimization method, the materials, surface roughness, and contact pressure of middle bearing and its sleeve can be optimized as shown in Table 2. To simulate the thermal behavior of the spindle, a finite element model was established as shown in Fig. 3. The three sets (W/(m2 K)) Materials of contact ball bearings were simplified to three rings, and the heat generated by bearing rotation can be calculated as follows. Qb \u00bc 1:047 10\u22124nM \u00f018\u00de where Qb is the heat generated by angular contact ball bearings (W); n is the spindle speed (rpm); and M is the friction torque (N mm) which can be calculated as follows. M \u00bc M 0 \u00feM 1 M 0 \u00bc 10\u22127 f 0 v0n\u00f0 \u00de2=3D3 m; v0n\u22652000 M 0 \u00bc 160 f 0D 3 m; v0n < 2000 M 1 \u00bc f 1P1Dm 8>< >: \u00f019\u00de where f0 is the coefficient related to the lubrication mode; v0 is the kinematic viscosity of the lubricant (mm2/s); f1 is the coefficient related to the load; P1 is the load applied to the bearing (N); andDm is the average diameter of angular contact ball bearing (mm)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002117_cp.2016.0128-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002117_cp.2016.0128-Figure3-1.png",
        "caption": "Figure 3: Bar with both ends fixed",
        "texts": [
            " = ( T) (1) in which is the coefficient of thermal expansion (CTE) with unit 1/K or 1/\u00b0C and T is the increase in temperature with unit K or \u00b0C. A positive means the object is expanded and a negative means the object is compressed. If a bar with length L is subjected to temperature increase and free to expand, the change in length due to the thermal expansion can be expressed by Equation (2). However, this deformation will not produce any stress and the bar is a statically determinate structure. (2) However, if the bar has both ends fixed as shown in Figure 3 and is exposed to uniform temperature increase T, reaction R will be developed over the bar and the bar will be subjected to compressive stresses. However, the thermal expansion in the transverse direction does not produce any stress due to no constrains or supports applied. Therefore, the problem can be simplified to a 1D problem in the axial direction. From another point of view, assuming that the fixed support at the right end in Figure 3 is removed and the same uniform temperature is applied, with force R applied from the right support pointing to the left, the bar will still keep its original length L in the axial direction. The displacement produced by force R on the bar is (3) where E is the modulus of elasticity or tensile modulus of the material with unit Pa and A is the cross section area of the bar. Besides, the displacement induced by thermal expansion is equal to the force R induced displacement as shown by Equation (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000991_s00170-015-6812-0-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000991_s00170-015-6812-0-Figure17-1.png",
        "caption": "Fig. 17 Three-dimensional modeling diagram of casting for vehicle and the main parts of die casting die. a Shell part (aluminum alloy brand: ADC12). b The shell part with injecting system and overflow. cMain parts of the die casting die: A moving die; B fixed die; C die casting; D, E, F, G movable cores",
        "texts": [
            "7 Microhardness The microhardness of the laser remelting area is shown in Fig. 16. The microhardness changes slightly for different laser parameters. The microhardness achieved in the laser remelting area is higher than that of the H13 matrix. The average microhardness of the melted zone and the transition zone is 620 and 580 HV, respectively. The microhardness of the matrix is 501 HV. In order to verify the effectiveness of the laser remelting process, an aluminum shell part for the vehicle is selected. Figure 17a shows the shell part made of the aluminum alloy ADC12, Fig. 17b the shell part with injecting system and overflow, and Fig. 17c the main parts of the corresponding die casting die which are made of the H13 die material. The weight of the shell part is 0.65 kg. The size of the shell is 155\u00d768\u00d734mm. The wall thickness is uneven, and the maximum and minimum wall thicknesses are 34.2 and 3.95 mm separately. The minimum radius of the shell is only 0.4 mm. The die is composed of the fixed die (A), the moving die (B), and four movable cores (D, E, F, G), which are made of the H13 die material. The size of the die in shut state is 350\u00d7220\u00d7170 mm. As the wall thickness is uneven and the radius is small, the service life of the die is only 13,000 shots. In real die casting production, defects on the outer surface of the casting appear firstly on the edges with a small radius, as shown in the red circles in Fig. 17a. According to the defects of the die casting die, the fixed die and the moving die were processed by laser remelting. The laser processing parameters are listed below: Electric current 150 A Frequency 5 Hz Pulse duration 8 ms Single-pulse energy 333 J Defocus distance 5 mm Scanning speed 0.5 mm/s The manufacturing procedure of the die is as follows: (1) rough milling; (2) quenching and operation; (3) rough EDM with allowance 0.15 mm, as shown in Fig. 18; (4) laser remelting process, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000528_978081000342.349-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000528_978081000342.349-Figure3-1.png",
        "caption": "Figure 3: Cut through the full floating ring bearing and free body diagram",
        "texts": [
            " Further on, the ring speed ratio of the floating ring and the rotor rotational speed can now be validated. The complexity of the model is intentionally kept low to show basic effects related to the journal bearing such as oil-whirl. Further, one can compare the numerical calculated critical threshold speeds with the measured ones. Increasing the modelling depth step by step, taking gyroscopic effects into account and using a finite element description of the rotor are scheduled for future investigations. Figure 3 left shows the axial view on the full floating ring bearing. The position of the rotor is given in Cartesian coordiantes and ; the position of the floating ring is given by and . The rotor and the floating ring are rotating with \u03a9 and \u03a9 , respectively. The later used transformation = + , \u2217 = + (1) into complex coordiantes (see [2] and [10]) with as the complex number, reduces the number of variables. For describing the full floating ring bearing, the following assumptions are introduced:  incompressible fluid,  constant and equal viscosities in both oil films,  constant pressure distribution in radial direction,  fully enclosed cylindrical bearings, no drill holes, no defects (e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002387_cae.21769-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002387_cae.21769-Figure3-1.png",
        "caption": "Figure 3 Quadcopter diagrams for movements along the vertical axis. (a) Hovering. (b) Ascending. (c) Descending. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com].",
        "texts": [
            " To move the quadrotor along the vertical axis, the speeds of the four motors are either increased to raise the quadcopter up or decreased to lower it down. To keep the quadcopter still in the air, the total thrust produced by the propellers must compensate for the earth\u2019s gravity and the torques produced by the four motors must sum to zero. The second condition is achieved by rotating opposite motors in the same direction while rotating neighboring motors in the opposite direction while maintaining the same angular speed for the four motors. Figure 3 shows diagrams illustrating quadcopter movements along the vertical axis, where the thickness of the arrows indicates the speed level of the rotors. The total thrust u produced by the four propellers will be calculated based on the following vertical movement modes: hovering, ascending and descending translations. This thrust u always points towards the positive zb axis and is given by Equation (8). u \u00bc X4 i\u00bc1 f i \u00bc k X4 i\u00bc1 vi 2 \u00f08\u00de Multiplying Equation (1) by vector ~u \u00bc 0 0 u\u00bd T gives the total thrust ~FR expressed with respect to the inertial reference frame as shown in Equation (9)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001299_romoco.2013.6614602-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001299_romoco.2013.6614602-Figure1-1.png",
        "caption": "Fig. 1. Configuration and kinematic parameters of the N-trailer robot",
        "texts": [
            ") : JR x JR f-7 JR is a continuous version of the four-quadrant function Atan2 (., .) : JR x JR f-7 (-7f,7fl . Decision factor CT E {-1,+1}, introduced in (5) and included in partial derivatives Fx and Fy, detennines a desired quadrant for reference orientation ed(X, y). Now, by referring to the last-trailer kinematics in the form of (2), one can introduce the path-following error for the last trailer e(qN)\ufffd [F(iiN)] \ufffd [ CTf(iiN2 ] EJR2, (7) eo(qN) eN - ed(qN) where iiN = [XN YNlT is a position vector of the guidance point P (cf. Fig. 1), and f(iiN) == f(XN,YN) determines function f(x,y) evaluated at point iiN. In definition (7) the first component F( iiN) can be called the signed distance value (see [12]) between the guidance point P and the desired path (note that F(iiN) = 0 only when the last trailer is on the desired path), while the second component, eo (q N ), is the orientation error where, according to (6), ed(iiN) = Atan2c (-Fx(iiN),Fy(iiN)) . The control problem under consideration is to find a bounded feedback control law uo(f3 , qN, t) which guarantees that error (7) is convergent in the sense: lim F(iiN(t)) = 0, lim eo(qN(t)) = 2v7f, v E Z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003329_aim43001.2020.9158972-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003329_aim43001.2020.9158972-Figure5-1.png",
        "caption": "Figure 5. Appearance of IMU sensor installed between the top of the robot and first unit",
        "texts": [
            " However, since a cumulative error occurs during the integration operation, a correction is performed by focusing on the periodic motion of the robot. Bending pipe discrimination is determined from the change in angular velocity and the attitude angle of the sensor. The angle of the bending pipe is specified by the standard, and the type is limited. The existence of a bending pipe is confirmed by focusing on the change in angular velocity. Then, the angle of the bending pipe is determined using the attitude angle of the sensor. A. Installation of the IMU on the robot The IMU sensor is housed in a protective case, as shown in Fig. 5. This protective case is fixed between the head part of the PEW-RO V and the first unit part using vinyl tape. The IMU sensor is a TAG250 (Tamagawa Seiki, Co., Ltd., acceleration range \u00b16 G, angular velocity range \u00b1200\u00b0/s). In addition, the IMU sensor is installed in the direction where the robot crawls, which is positive on the y-axis. The IMU sensor is connected to a PC with a USB cable that passes inside the robot body. It is possible to calculate the acceleration, angular velocity, and three-dimensional attitude angle of the sensor from the IMU using included software (Tamagawa Seiki Co"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003758_electronica50406.2020.9305143-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003758_electronica50406.2020.9305143-Figure1-1.png",
        "caption": "Fig. 1. The structure of a sample roller bearing",
        "texts": [
            " The largest share belongs to the bearings (41%), which according to IEEE studies, have been extracted on engines above 200 hp and then the stators (27%) and the error related to the rotors (10%) [4, 5]. Therefore, in the actuators, which the PMSM motors are the central part, the bearing faults are more reported with respect to the other kind of faults. In [6,7], the bearing faults of the reaction wheel as the primary actuator in the attitude control subsystem are inspected. They used the model-based and signal-based methods for FDIR porpuses. In this paper, the rolling types of bearings are studied due to the greater use of these type bearings in PMSMs which is shown in Fig. 1. In general, faults related to roller bearings can be divided to the outer-ring, inner-ring, lubricant loss, and imperfect ball according to its mechanical characteristics. The defects related to each part have different effects on the measured current signal of each phase on the motor driver. As mentioned, the faults associated with each part of the bearing are related to a frequency [8,9], which, for example, the perforation of the hole on the outer ring or the inner ring, can evoke a fault related to them"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000801_tasc.2012.2231953-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000801_tasc.2012.2231953-Figure2-1.png",
        "caption": "Fig. 2. SMB structures of (a) type-1 and (b) type-2.",
        "texts": [
            " The levitated table moves following the conveying table due to flux pinning forces. Signals of table displacement are put into a personal computer (PC) through A/D converters. Control signals from D/A converters are put into power amplifiers. Then, left and right coils are excited to suppress the table vibrations by using PD (proportional and derivative) control. In magnetic levitations, levitation forces are affected by the magnetic field in the radial direction. Then, two types of SMBs with a central superconductor are studied as shown in Fig. 2. Fig. 2(a) shows a SMB composed of a doughnutshaped superconductor (\u2248 \u03c649 mm \u00d7 \u03c624 mm \u00d7 25 mm), a cylindrical superconductor (\u03c69.0 mm \u00d7 25 mm) in the center of the 1051-8223/$31.00 \u00a9 2013 IEEE doughnut-shaped superconductor, a levitated disk-shaped PM (\u03c630 mm \u00d7 5 mm, surface magnetic flux density: 0.3 T) and an electromagnet (EM) under the disk shaped superconductor, corresponding to type-1 SMB. The cylindrical superconductor is used for levitation and for making magnetic flux path between two superconductors. The levitated PM in Fig. 2(b) is a doughnut-shaped PM (\u03c630 mm \u00d7 14 mm \u00d7 5 mm, surface magnetic flux density: 0.3 T), corresponding to type-2 SMB. Table I shows the lists of these parameters. Fig. 3 shows the magnetic flux density distributions in the radial direction of the type-1 and type-2 SMBs. The magnetic flux density is measured on the surface of PMs. In the case of type-1 SMB, magnetic flux density distribution shows a point symmetry around the center of PM. In the case of type-2 SMB, there is a doughnut- shaped magnetic flux distribution as shown in Fig. 3(b). The gradient of magnetic flux density is expected to improve the levitation force of type-2 SMB shown in Fig. 2(b). Repulsion forces between HTSC and PM are measured. In the measurement, a z-stage and a load cell are connected to the levitated PM. The attraction and repulsion forces are measured by using the load cell for various applied currents from \u22121.0 A to 1.0 A. Fig. 4 shows the relationships between attraction/ repulsion forces and applied current, representing the force (a) in the horizontal direction and (b) in the vertical direction. The repulsion force and attraction force show minus sign and plus sign, respectively",
            " The experimental results show that the damping coefficient of type-2 SMB is larger than that of type-1 SMB. This is because the gradient of magnetic flux density in Fig. 3(b) is larger than that in Fig. 3(a). That is, the type-2 SMB is more effective than the type-1 SMB. Hereafter, the type-2 SMB is adopted for the magnetically levitated conveyer in this paper. NAKAYA et al.: DYNAMIC CHARACTERISTICS OF MAGNETICALLY LEVITATED CONVEYER USING HIGH Tc SMB 3601304 B. Impulse Response of Two-Axis Control Our group applied the type-2 SMB [Fig. 2(b)] to the magnetically levitated conveyer. Then, the levitated table is supported by two type-2 SMBs. Impulse responses of the magnetically levitated table are studied. In the experiments, impulses are applied to the levitated table in the vertical direction. Then, the displacement of the levitated table in the vertical direction is measured. The displacement is measured by using a laser displacement sensor. Fig. 6 shows the experimental result of impulse responses (a) without control and (b) with control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure5-1.png",
        "caption": "Fig. 5. Assembling stator of a novel homopolar-type MB unifying the four C-shaped cores. (a) Assembling two stator cores and a case. (b) Assembling stator cores and four winding cores.",
        "texts": [
            " In addition, the width between the magnetic poles, shown in Fig. 4(b), easily becomes small by the unified stator laminated cores. It is noted that the leakage flux between the magnetic poles is blocked by inserting small magnets between the magnetic poles, as shown in Fig. 4(b). Since the novel homopolar-type MB has no large magnets for generating bias flux, the magnet volume compared with that in the general homopolar-type can be decreased. Thus, the cost of the novel homopolar-type MB can be reduced. Fig. 5 shows the procedure of a stator assembly for the novel homopolar-type MB in detail. The stator-laminated cores are inserted into a nonmagnetic material case from the axial direction, as shown in Fig. 5(a). In addition, small magnets are inserted to the holes of the stator-laminated cores. Subsequently, it is possible to assemble the stator easily by inserting the stator bulk cores with the windings to the predefined position of the stator-laminated cores, as shown in Fig. 5(b). Accordingly, the stator inner surface accuracy is high because of the unified stator laminated cores. Fig. 6 shows the principle of generating the suspension force Fb+ toward the positive direction in the b-axis. The x-axis and the y-axis are defined as the perpendicular coordinates fixed to the stator with reference to the winding position. In addition, the a-axis and the b-axis are converted from the x-axis and the y-axis using rotational conversion with 45\u00b0 toward the clockwise direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003833_b978-0-12-800054-0.00026-5-Figure26.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003833_b978-0-12-800054-0.00026-5-Figure26.1-1.png",
        "caption": "Figure 26.1 PWA F-100 engine compression rigid stator link produced by press and sinter of BE powder.",
        "texts": [
            " The conference was well attended by key university and manufacturing company technologists involved in advancing the state of the art in powder metallurgy. From that conference, a number of Air Force\u2013sponsored programs emerged that advanced the application potential for titanium and superalloy PM technologies in both engines and airframes. The program with PWA noted above performed a program with powder supplier REM Metals together with press and sinter consolidation sources Gould Laboratories and Imperial Clevite to establish a process for the production of a near-net F100 engine compression rigid stator connecting link, Figure 26.1. The process produced parts with 91% theoretical density with static strength properties comparable to wrought material but with somewhat lower fatigue properties that were judged adequate for the intended application. Economic analysis of the process showed that an acquisition cost reduction of 71% was possible in production lot quantities over the then current bill of material component that was machined from bar stock. It was recommended that the process be considered for application to the engine and that press and sinter technology be applied to other engine parts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002635_ecce.2016.7854747-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002635_ecce.2016.7854747-Figure17-1.png",
        "caption": "Fig. 17. The 6/7-pole SCPMM prototype. (a) Stator. (b) Rotor.",
        "texts": [],
        "surrounding_texts": [
            "As a pivotal feature for the wide-speed-range applications, the flux-weakening performance of 6/7-pole machine should be analyzed. First, a flux-weakening coefficient is utilized to quantify the capability to extend the constant-power operating range, which can be defined as ( ) ( ) max max, 0 = 0 d d q fw PM q L i I i I k i\u03c8 = \u2212 = = (21) where id and iq are the d/q-axis currents. Ld and PM are the daxis inductance and the PM flux linkage with consideration of cross-coupling effect [18]. Imax is the maximum phase current. The corresponding flux-weakening characteristic parameters are listed in Table IV. It can be observed that the maximum achievable speed range can be expanded via reducing the PM flux linkage owing to the increase of flux-weakening factor. The efficiency maps under the full and weak magnetization states are shown in Fig. 16. It can be seen that the efficiency improvement over a wide operating regions can be achieved since the highest efficiency region locates at different speed regions for different magnetization levels. This is mainly attributed to the significant iron loss reduction at the weak magnetization state. For example, the higher magnetization level is preferred for constant-torque operation to obtain high torque output and high efficiency, whereas the weak magnetization state is favorite to the constant-power operation region with reduced iron losses, which are the dominant loss component at high-speed region. VI. EXPERIMENTAL VERIFICATION In order to validate the previous FE analysis, a 6/7-pole SCPMM prototype is manufactured and tested. The stator and rotor assemblies of the prototype are shown in Figs. 17(a) and (b), respectively. Fig. 18 shows the FE predicted and measured open-circuit back-EMFs under different magnetization states under 400r/min. The magnetization control board is characterized by a capacitor charging/discharging based circuit as detailed in [10], which allows a maximum magnetizing current reaching up to 100A. It can be seen that the FE predictions agree well with the measurements. The manufacturing tolerance and end effect are mainly responsible for the slight discrepancy. Meanwhile, the predicted and measured root-mean-square EMFs as functions of the current pulse magnitudes are shown in Fig. 18(b), in which the satisfactory agreement confirms the excellent flux regulation capability of the proposed design. It can be seen that the remagnetization control is relatively harder than demagnetization, and the fully remagnetizing current amplitude is about 3 times over the fully demagnetizing one. Meanwhile, the RMS back-EMFs drop linearly with increasing the applied demagnetizing current, while the remagnetization curve experiences a significant non-linearity. This is mainly attributed to the nonlinear B\u2013H hysteresis feature of AlNiCo PM. Afterwards, the torques as functions of the q-axis current at the full magnetization state is measured. As shown in Fig. 19, the good agreement between the measured and FE predicted torques can be obtained. VII. CONCLUSIONS In this paper, a novel SCPMM is proposed, in which the LCF AlNiCo PMs are alternately placed between the adjacent stator teeth. The proposed machine benefits from simplified online PM magnetization, good demagnetization withstand capability, robust rotor and easy thermal management. Meanwhile, the excellent flux regulation can be achieved with negligible excitation copper loss. A 6/7-pole prototype is manufactured and tested to validate the FE analysis experimentally. Some key findings are summarized as follows: 1) The proposed machines operate based on the variable reluctance behavior between stator and rotor slient poles. Thus, the bipolar coil flux linakge can be obtained. Meanwhile, the new machines behave similarly as magnetically geared machines, in which plentiful modulated air-gap field harmonis contribute to torque production. 2) The machines having even rotor poles suffer from significant even-order EMF harmonics and large torque ripple. On the other hand, the 6/7-pole machine exhibits symmetrical back-EMF and excellent torque quality, which benefits from the cancellation of even-order flux-linakge harmonics between two individual coils in one phase. 3) The high efficiency operation can be extended with a wide range of speeds and loads by combining the highest efficiency characteristics under each magnetization state."
        ]
    },
    {
        "image_filename": "designv11_30_0003759_ssci47803.2020.9308272-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003759_ssci47803.2020.9308272-Figure1-1.png",
        "caption": "Fig. 1: The three modules that compose our Robots",
        "texts": [
            " The behavioral ones describe properties only measurable during the lifetime of the robot, such as speed or balance. Using such traits we will inspect the evolved populations while addressing the following specific sub-questions: \u2022 Which morphological and behavioral traits differ the most between the two setups? \u2022 In which way do morphological and behavioral traits differ in the two setups? Our modular robot system is based on RoboGen [3]. Each robot is composed of three different types of modules: one Core module (Figure 1a), an arbitrary number of Brick modules (Figure 1b), and an arbitrary number of Joint modules (Figure 1c). The Core module is unique for each robot and represents the robot \u201chead\u201d that, in the original physical incarnation [8], contains the main logic board and the battery. The Core module has four connection points where other modules can be attached. Brick modules represent the \u201cbackbone\u201d of the robot. Only through Brick modules, the robot can take up arbitrary shapes. Actuation can only be achieved through the Joint modules, thus Joint modules are the only modules capable of changing the state of the robot in the environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001270_iciea.2013.6566330-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001270_iciea.2013.6566330-Figure1-1.png",
        "caption": "Fig. 1. A typical sketch of a VTOL aircraft in the vertical plane [1].",
        "texts": [
            "ndex Terms\u2014Vertical take-off and landing aircraft (VTOL), second order sliding mode (SOSM), adaptive sliding surface. I. INTRODUCTION The vertical take-off and landing (VTOL) aircraft is a highly complex nonlinear system whose aerodynamic parameters vary considerably during the flight. Real world examples of the VTOL aircraft system are aerospace vehicles like helicopters, rockets, balloons and harrier jets. All aerospace vehicles are difficult to model due to their changing aerodynamic parameters and environmental behavior during flight. Fig.1 shows the typical coordinate system for a VTOL aircraft in the vertical plane. VTOL aircraft system is a an uncertain system which is affected by both matched and mismatched types of uncertainties. For controlling such uncertain systems, the sliding mode controller (SMC) [2] has been applied widely because of its simplicity and inherent robustness. However, chattering in the control input is an undesired phenomenon in the conventional first order SMCs. Furthermore, the design prerequisite of advance knowledge about the upper bound of the uncertainty is a stringent condition in the case of conventional SMCs and is often difficult to meet in practice"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.60-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.60-1.png",
        "caption": "Figure 3.60 Worm gearhead schematic",
        "texts": [
            " Planetary gearheads can be fabricated with one, two or three passes in tandem providing ratios from 3:1 to 500:1 with efficiencies of 85 to 96%, in sizes 14 to 160. 110 Electromechanical Motion Systems: Design and Simulation A hybrid gearhead consists of one or two input spur gear stages; with a planetary gear output stage. This arrangement can provide a high ratio design that is lower in cost than a multi-pass planetary design. It is often used in right angle gearheads in which the input pass is a spur gear mesh followed by a right angle bevel pass that drives the output planetary gearing. A worm gear assembly consists of two components, as shown in Figure 3.60. The input is the worm shaft, similar to a lead screw. The output is obtained from the worm gear that meshes with the worm. This arrangement is unidirectional in that the worm cannot be driven by the worm gear and actually will bind and result in a braking action if the load attempts to back drive. Worm gearing has traditionally been used for high power, unirotational applications to obtain right angle, high reduction ratios in a compact assembly in which efficiency is of secondary concern. Contemporary improvements in tooth design, materials, lubrication, bearings and thermal design have created worm gear assemblies that are compatible with the severe requirements of closed loop servo operation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000712_itsc.2015.311-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000712_itsc.2015.311-Figure7-1.png",
        "caption": "Fig. 7. a) 3D configuration space; b) Discretized orientations; c) Motion primitives end in different orientations",
        "texts": [
            " 6) Pnorm : nominative path dist(, ) : function, which gets distance between two points 1 Function ret GetClearanceMap(Pnorm) 2 initiates all occupied in path clear with 0; 3 foreach pi \u2208 Pnorm do 4 foreach cell \u03b6i, dist(\u03b6i, pi) < clear[pi] do 5 clearance\u2190 path clear[\u03b6i]\u2212 dist(\u03b6i, pi); 6 if clearance > path clear[\u03b6i] then 7 path clear[\u03b6i]\u2190 clearance; 8 return path clear[]; A virtual active window is assumed to be attached with the mobile robot and local path planning is constrained inside the window. Only the dynamic obstacles inside the window are considered. The active window centers on the robot and moves along with it. A motion primitives-based search is applied inside the active window. The search uses a set of motion primitives in 3D configuration space: C = {(x, y, \u03b8)|0 \u2264 \u03b8 < 2\u03c0} (Fig. 7(a)). Orientation \u03b8 is discretized into 32 entries [as in Fig. 7(b)]. Fig. 7(c) is a simplified illustration of how the motion primitives are applied beginning from node ni in a certain orientation \u03b8i, and ending at nodes in neighboring orientations. As in Fig. 8, the green frame is the active window attached to the robot; motion primitive-based path planning is applied locally inside both the active window and the GPC. The blue object represents a dynamic obstacles that the robot has encountered on the way. An evaluation of the nodes are based on the A* algorithm [21], which includes two terms [as shown in (1)]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure1-1.png",
        "caption": "Fig. 1.(a) a direct-through hollow-shaft cooling structure, (b) a recirculating hollow-shaft cooling structure.",
        "texts": [
            " Rotor is usually more difficult to cool than stator, especially in vacuum applications, because the airgap around the rotor is equivalent to insulating material. To enhance the heat transfer, forced oil cooling is widely used because oil has a higher thermal conductivity and density than air. The way that the motor stator and the rotor are immersed in oil can sufficiently cool the motor [9]. However, significant friction loss may occur on the rotor surface due to the rotor rotation. In order to reduce the friction loss, a hollow-shaft cooling method can be used [10]-[11]. As illustrated in Fig. 1(a), the oil is forced to pass through a direct-through hollow-shaft by two rotary couplings. Yaohui Gai et al. investigate the CHTC of the direct-through hollow-shaft in [12] and [13]. The inlet and outlet of the direct-through hollow-shaft are located at both ends of the motor, and the installation of load will be inconvenient. Fig. 1(b) presents a recirculating hollow-shaft cooling structure for high speed PMSMs. The number of dynamic seals in the recirculating hollow-shaft is half of the direct-through hollow-shaft, and the inlet and outlet are on the same side. Compared with the direct-through hollow-shaft cooling structure, its coolant is introduced via a stationary inner tube and is fed back to the front due to the function of tail baffle [14]. However, recirculating hollow-shaft cooling structure has not been studied for its CHTC and friction loss caused by high speed rotation on the rotating surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003248_j.cja.2020.06.030-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003248_j.cja.2020.06.030-Figure2-1.png",
        "caption": "Fig. 2 Relief grinding coordinate system of right-handed hob.",
        "texts": [
            " In order to analyze the radical relief grinding movement of grinding wheels correspond to gear hobs, the model based on the spatial motion relationship of the above two objects in an arbitrary moment of the machining process is constructed in Fig. 1. where SH XH;YH;ZH\u00f0 \u00de and SW XW;YW;ZW\u00f0 \u00de are respectively the moving coordinate systems fixed to the gear hob and grinding wheel. ZH and ZW are respectively the axes coincides with hob shaft and grinding wheel shaft. OH and OW are the axis midpoints of the gear hob and grinding wheel. The key to calculating the grinding wheel axial profile is the analysis on relative location of two objects. As shown in Fig. 2, the workpiece that needs to be ground is a right-handed gear hob and in Fig. 3, it is an annular buckle gear hob. S X;Y;Z\u00f0 \u00de in Figs. 2 and 3 is a fixed spatial coordinate system; k is the angle between the gear hob and grinding wheel rotation axes, generally taking the helix angle on hob standard pitch circle and also regarding as the vertical installation angle of grinding wheel; h represents the relative turning motion parameter; A represents the distance between two rotation axes (ZH and ZW)",
            " For the right-handed or left-handed hob, it performs a helical motion along its own axis, and the helical motion parameter P can be generally written with the following equation P \u00bc Pn 2p \u00f01\u00de where Pn is the hob normal pitch, shown as follows: Pn \u00bc pmnn \u00f02\u00de rinding wheel axial profiles based on gear hobs, Chin J Aeronaut (2020), https:// where mn and n are respectively hob normal modules and thread number. As for the annular buckle gear hob, the value of helix angle k is equal to zero, and accordingly, there is no helical motion in hob, neither the helical motion parameter P. When the hob rotates per unit angle, radical movement distance of the grinding wheel along the gear hob can be gotten KP \u00bc K nZK 2p \u00f03\u00de where ZK is the number of chip-holding groove; K is the reliving quantity of gear hob. 2.2. Coordinates conversion It is assumed that point Q shown in Fig. 2 is an arbitrary engaging point between grinding wheel and gear hob in the grinding process. To achieve the transformation relation of point Q in coordinates, transition should be made from the Please cite this article in press as: YANG S, CHEN W Modeling and experiment of g doi.org/10.1016/j.cja.2020.06.030 gear hob coordinate system to the grinding wheel coordinate system. In Fig. 2, three reference coordinate systems are presented, supposing the unitary vectors of S X;Y;Z\u00f0 \u00de along the X,Y and Z axes are i, j and k, respectively. It can be assumed, as above, that the unitary vectors of SH XH;YH;ZH\u00f0 \u00de along the XH,YH and ZH axes are iH, jH and kH, the unitary vectors of SW XW;YW;ZW\u00f0 \u00de along the XW,YW and ZW axes are iW, jW and kW. The position vectors of the point Q relative to the origin of coordinates O, OH and OW are respectively r \u00bc \u00bdx; y; z T, rH \u00bc \u00bdxH; yH; zH T and rW \u00bc \u00bdxW; yW; zW T",
            " The transformation relation from coordinate system SH XH;YH;ZH\u00f0 \u00de to SW XW;YW;ZW\u00f0 \u00de can be described as the following equations r \u00bc Mh ZrH \u00fe Phk \u00f04\u00de rW \u00bc Mk Xr A KPh\u00f0 \u00deiW \u00f05\u00de where Mh Z and Mk X are coordinate transformation matrices that rotate around Z axis in a clockwise direction and around X axis in a anticlockwise direction, respectively. The matrices are denoted as Mh Z \u00bc cosh sinh 0 sinh cosh 0 0 0 1 2 64 3 75 \u00f06\u00de Mk X \u00bc 1 0 0 0 cosk sink 0 sink cosk 2 64 3 75 \u00f07\u00de Then, coordinate values of the meshing points in SW XW;YW;ZW\u00f0 \u00de of Fig. 2 can be determined with the equations below: xW \u00bc xHcosh yHsinh\u00fe KPh yW \u00bc xHsinh\u00fe yHcosh\u00f0 \u00decosk\u00fe zH \u00fe Ph\u00f0 \u00desink zW \u00bc xHsinh\u00fe yHcosh\u00f0 \u00desink\u00fe zH \u00fe Ph\u00f0 \u00decosk 8>< >: \u00f08\u00de In the case where the grinding object of the grinding wheel is an annular buckle gear hob, shown in Fig. 3, it is only necessary to set the value of P to zero, therefore, coordinate values of the meshing points can be reduced to the following form: xW \u00bc xHcosh yHsinh\u00fe KPh yW \u00bc xHsinh\u00fe yHcosh\u00f0 \u00decosk\u00fe zHsink zW \u00bc xHsinh\u00fe yHcosh\u00f0 \u00desink\u00fe zHcosk 8>< >: \u00f09\u00de The key point of setting the value of P to zero also applies to the following conditions when it comes to annular buckle gear hob. The next calculations take the gear hobs with hand of helix as examples. In the coordinate system SH XH;YH;ZH\u00f0 \u00de of Fig. 2, the shaping motion of gear hob is comprised of three parts. The first part is hob rotation, the second is translational motion of the hob along its axis (ZH) and the last is grinding wheel radical feeding rinding wheel axial profiles based on gear hobs, Chin J Aeronaut (2020), https:// movement. The grinding wheel rotation can be omitted since it is not related to the helicoid-formation. Assuming that gear hob angular velocity is xH, the linear velocity of meshing point Q with the gear hob movement can be determined from the following equation: mH \u00bc xH kH rH \u00fe PkH\u00f0 \u00de \u00f010\u00de and the following denotation on linear velocity of meshing point Q with grinding wheel movement is adopted mW \u00bc KPiW \u00f011\u00de According to the Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001244_aim.2014.6878110-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001244_aim.2014.6878110-Figure1-1.png",
        "caption": "Fig. 1. Swimmer with two links, i.e., helical tail and spherical body with local Frenet-Serret frames (bnt) along the tail, swimmer frame (sqr) on the center of mass, and lab frame (xyz).",
        "texts": [
            " We imposed the effective elasticity of the surface contact and demonstrated that, once the channel is implemented by means of contact force based on penalty method and appropriate frame rotations, swimmer follows the cylindrical wall exhibiting the influence of gravitational pull and lubrication under different environmental conditions with great numerical stability. II. METHODOLOGY RFT-based model studied in this paper is already discussed in great detail in [12], [27], [28]; however, the equation of motion will be discussed for the sake of introducing asymptotical and numerical corrections to the resistance matrices of the body and tail (see Fig. 1). The RFT method is valid for viscous flows and creeping motions in small scales. In such swimming conditions, inertial accelerations in hydrodynamic and robotic equations are all neglected due to the comparisons on order of magnitude. The final equation of motion dictates that the sum of overall hydrodynamic resistance of robots own motion and channel flow, gravitational pull, and contact forces, i.e., \u03a3F, acting on the links of the swimmer should add up to zero [29]: { } { } { }b,t b t drag contact gravity b t drag contact gravity b,t b,t 0, 0, . F F F F F T T T T T F V B T \u2126 + + + + = + + + + =        = \u2212       (1) Here, B stands for the resistance matrix of the body and tail, which are denoted by subscripts \u2018b\u2019 and \u2018t\u2019, respectively (see Fig.1), and vectors V and \u2126 are the surface velocities of swimmer [27]. The swimmer geometry selected in this study is that of an E. Coli minicell, i.e., a single-celled organism with one helical flagellum and one spherical body rotating in opposite directions [30] due to the motor torque principle [22]. Fdrag denotes the hydrodynamic drag exerted on the swimmer by the upstream, i.e., channel flow, and Fcontact denotes the structural reaction force along the surface normal of the channel at the point of contact (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002601_humanoids.2016.7803410-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002601_humanoids.2016.7803410-Figure3-1.png",
        "caption": "Fig. 3: IK solution for putting the box on the table.(Each subfigure includes a robot pose and its CoM/support polygon diagram. The distance between the table and the robot in each of these subfigures is different: 0.6m in (a)-(b), 0.8m in (c)-(d), and 1.0m in (e)-(f). (a) shows the initial pose of the robot. (b)-(f) are the IK solutions.)",
        "texts": [
            " Therefore, to generate a feasible IK solution for Valkyrie, a variety of costs and constraints need to be set, such as, joint displacement costs using a normal standing pose as the nominal pose, pelvis height and orientation costs, torso orientation costs, joint limit constraints, collision avoidance constraints, and CoM constraints. According to the task requirements and different foot placement strategies, some specific costs and constraints also need to be added, which are explained by a box placing example below. The task constraints are set based on the task requirements. Most of them introduce the interaction between the end effector and the environment. For example, in Fig. 3a, Valkyrie is standing in front of a table and attempts to place the box on it. The task requirements are moving the hand on top of the table and keeping both hands static relative to each other to hold the box. Therefore, two task constraints need to be applied. The first one is a first type Cartesian posture constraint on the left hand. The desired left hand posture ddesired is provided according to the position of the table (see Fig. 3b). The second one is also a first type Cartesian posture constraint which asks the right hand to keep relatively static with respect to the left hand. The desired right hand posture can be computed by multiplying the current transform between the left hand and the right hand with the desired left hand posture. The box is grouped into the robot model as calculating the collision avoidance constraint. When the robot is close to the table, as it in Fig. 3b, it can easily place the box on the table while keeping the CoM in the support polygon without moving the feet. However, after moving the table 20 cm to the left, as shown in Fig. 3c, the IK solution with both feet standing in the same position violates the CoM constraint. By relaxing the constraint on one foot, the support polygon can be reshaped in order to position the CoM inside the polygon (see Fig. 3d). When the table is moved to the left 40 cm, although the robot can meet the CoM and end effector target constraints shown in Fig. 3e, many joints exceed their torque limits, such as the back and hip joints. To deal with the problem, the robot can be allowed to move both of its feet. The IK solution can not only provide final joint states but also a desired standing position (see Fig. 3f). Thus, these three types of foot position strategies, such as fixing both feet, fixing one foot and relaxing both feet, require three different constraint setups. 1) Fixing both feet: Add the first type Cartesian posture constraints on both feet. The desired postures of both feet are their current poses. 2) Fixing one foot: Add a second type Cartesian posture constraint for the movable foot, whose details are described in Section II-A. Add a first type Cartesian posture constraint on the stationary foot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002634_edpc.2016.7851310-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002634_edpc.2016.7851310-Figure2-1.png",
        "caption": "Fig. 2 Lamination stack with magnet fixation based on adhesives",
        "texts": [
            " Thus, a concept of evaluating imbalances was developed using 2- and 3-dimensional vector combination. In addition, joining tests were conducted to assess the achievable radial runout. II. CONCEPT OF SELECTIVE BALANCING The developed concept is based on current manufacturing technologies, particularly regarding permanent magnet synchronous drives with embedded magnets (IPMSM). Rotors of these highly efficient and powerful drives are usually assembled using a set of preassembled lamination stacks with magnets. As shown in Fig. 2, the magnets are fixed with adhesive. 978-1-5090-2909-9/16/$31.00 \u00a92016 IEEE Magnet assembly and the use of adhesives create an arrangement of uneven distributed auxiliary materials, amplifying existing imbalances within the lamination stack. Since balancing is usually executed after finishing the complete rotor assembly, assuming a worst-case scenario, these imbalances may add up to a significant amplification of existing 1st-order imbalances, Fig. 3 a). By considering the various values and directions of single imbalances and an assembly sensitive to these values, the resulting imbalance of the entire assembly can be reduced significantly, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure12-1.png",
        "caption": "Figure 12. Terminology for cylindrical cantilever of radius r in tension. Force F is applied tangent to the surface at the end of the beam.",
        "texts": [
            " For technological materials, G can often be calculated from E, but for biological materials\u2014non-Hookean, anisotropic\u2014G (or GJ) must be measured separately. Standard formulas exist for calculating J for simple shapes (e.g., for a solid cylinder: J=\u03c0r4/2) but biological structures are rarely simple enough to apply such equations. Thus, in practice, researchers usually measure the composite variable GJ, similarly to measuring EI. In this case, one end of the sample must be fixed so that a moment can be applied to the other end (figure 12). (For biological materials, this may require being glued into a holder with epoxy or cyanoacrylate cement.) The torsional stiffness is then given by: GJ ML , 11 q = ( ) where \u03b8 is the angular deflection in radians andM is the applied moment or torque (for a force F tangent to the surface of a circular cylinder of radius r, M=Fr). If a beam is loaded in pure compression, then by convention it is referred to as a column. Short, stubby columns being compressed act much like tensile structures in reverse: the structure shortens with a stress\u2013strain relationship that may have the same or very similar E, and they fail by crushing or by cracks allowing shearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001244_aim.2014.6878110-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001244_aim.2014.6878110-Figure6-1.png",
        "caption": "Fig. 6. Slightly heavier swimmer moving with lubrication effects and upstream in the channel; yz-trajectory of center of mass is obtained for 7500 complete cycles. Blue lines denote the channel walls.",
        "texts": [
            " 4; however, as the propulsive force pushes the swimmer further into collusion, the gait is changed and swimmer tends to follow the surface with a counterclockwise path once again. It is noted that, although not clearly shown in Fig. 5, swimmer slightly separates from the surface towards the end of the simulated time interval. Finally, upstream velocity and gravitational pull are introduced in the model. First, a small deviation in the density of the robot, e.g., 10% difference with the surrounding liquid, is investigated. Swimming robot, initially resting at the center, is actuated with the same rotational frequency and then follows the path presented in Fig. 6. In this scenario, the upstream velocity has a parabolic profile [14] having a maximum of 2 \u00b5m/s at the center. When swimmer reaches the 2rb barrier, the lubrication effect compels the robot to change its gait and move further down following the surface curvature. As the gravitational pull and propulsion force push the gait further down to the channel wall, contact force dominates the path akin to the previous scenario and swimmer starts to follow the surface with a counterclockwise rotation. Final scenario deals with an artificial swimmer made out of heavy materials; such as with an overall density seven times that of the surrounding liquid, which is an exaggerated scenario for some nano-scale magnetic artificial swimmers [35]; however, presents a suitable example on how surface contact would behave under impact due to high weight"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure7-1.png",
        "caption": "Fig. 7. Family of evolutoids for CW and CCW rotation of the input gear element.",
        "texts": [
            " 6 are two cases, CW rotation of the pinion element and CCW rotation of the pinion element. Viewing the involutoids from the end, the involutoids appear to have a singularity or cusp where its derivative is undefined. This is not the case as seen in the isometric views. The distance to the central axis is minimum and these values are identified as extremes. In the two cases presented, the transverse curve is identical. The family evolutoids define a surface, the base surface (family of evolutoids 3 ). Images of the base surfaces for CW and CCW rotation are presented in Fig. 7 . Base surfaces can be entirely on either side of the pitch surface (for different directions of rotation) or intersect the pitch surface as illustrated in Fig. 7 . Associated with each evolutoid is its involutoid curve. A segment of each involutoid can be used to define conjugate tooth flanks. Fig. 8 shows the tooth flanks relative to the base 3 One option is to identify the family of evolutoids as an evolutode . Continuing, the active tooth flank or family of involutoids becomes an involutode . Also, we will have a CW-evolutode and CW-involutode together with a CCW-evolutode and CCW-involutode. surfaces of Fig. 7 . Be careful to note that the active tooth flank is viewed from the toe end for CW pinion rotation whereas the active tooth flank is viewed from the heel end for CCW rotation. It was mentioned earlier that planar involute and evolute curves are partner curves where one curve can be defined in terms of the its partner curve. The evolutoid was first defined in terms of a transverse curve together with the pressure angle and spiral angle for this transverse curve. Subsequently, an involutoid curve was defined in terms of this evolutoid curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001740_s1560354715040024-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001740_s1560354715040024-Figure1-1.png",
        "caption": "Fig. 1. Laboratory and body-fixed coordinate systems.",
        "texts": [
            " It is assumed that the points are located on a horizontal plane passing through the body\u2019s center of mass and that the system\u2019s center of mass does not move inside the body. In contrast to [7], in the present work the local friction of rest is applied. 2. MODEL OF THE SYSTEM Consider a rigid body of mass m0, which is a hollow parallelepiped composed of six homogeneous rectangles. The body has length a, width b, and height 2h. The body rests on a rough horizontal surface. We introduce a laboratory coordinate system O\u2032xyz with origin on the surface and a coordinate system O\u03be\u03b7\u03b6 attached to the body, where O is the body\u2019s center of mass (Fig. 1). The axes O\u2032z and O\u03b6 are directed vertically upward, the axis O\u03be is parallel to the larger edge of the parallelepiped base, the axis O\u03b7 is directed so that the system forms a right-hand triple. We suppose that at the initial instant the corresponding axes coincide. The principal axes of the central ellipsoid of inertia of the body coincide with the coordinate system O\u03be\u03b7\u03b6. The moment of inertia with respect to the axis O\u03b6 is equal to C. The material points inside the body are located in such a way that the center of mass of the system coincides with the center of mass of the body O"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001204_roman.2013.6628553-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001204_roman.2013.6628553-Figure4-1.png",
        "caption": "Fig. 4. Spherical joint at the shoulder.",
        "texts": [
            " The other component of the cost function is the gravitational potential energy denoted by \u03b2 \u2211 mig T r\u03040 i . By adjusting the weighting factors \u03b1 and \u03b2, the minimization of the objective function will yield different results. For example, a large \u03b2 value suggests that in the final posture, the elbow should be kept low. A large \u03b1 value, however, usually means that the joint movements will be as little as possible. As discussed in the earlier Subsection, the position of the elbow is defined by joints q1 and q2. As illustrated in Figure 4, let (ex, ey, ez) be the the elbow position with respect to a local coordinate frame (Ox, Oy , Oz ) located at the shoulder joint. By geometry, it appears that the joint angles q1 and q2 can be computed as follows: ( ) Referring to equation (5)-(6), it should be noted that small variations in the task variables when ex and ez are close to zero, will lead to large variations in the joint variables. As a result, there is movement instability and this phenomenon is widely known as the kinematic singularity problem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001088_s11044-014-9417-8-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001088_s11044-014-9417-8-Figure4-1.png",
        "caption": "Fig. 4 (Left) Parallel mechanism consisting of serial chains with loop closures. Two base joints, q1, q3, as well as the elbow joints, q2, q4, and q6, are actuated while the remaining joints are passive. (Right) The position of the platform is commanded to move to a target while its orientation is uncontrolled. In this case, n = 9, mC = 6, p = 3, m = 2, N = 1, and k = 5",
        "texts": [
            " (40) Thus, (39) and (35) can be represented as the following system of 2n \u2212 k equations: ( 1 T ST u Sp 0 )( \u03c4 \u03bbu ) = ( h(q, q\u0307) 0 ) . (41) Given the inverse ( 1 T ST u Sp 0 )\u22121 = ( 1 \u2212 T ST u (Sp T ST u )\u22121Sp T ST u (Sp T ST u )\u22121 (Sp T ST u )\u22121Sp \u2212(Sp T ST u )\u22121 ) , (42) we have the following solution for the control torque: \u03c4 = ( 1 \u2212 T ST u ( Sp T ST u )\u22121 Sp ) h(q, q\u0307). (43) A block diagram of this control scheme is shown in Fig. 3. As an illustrative example of this control scheme, we consider the parallel mechanism depicted in Fig. 4(left), where n = 9, mC = 6, and p = 3. Units for this and all subsequent examples will be expressed in SI units (meters, seconds, and Newtons). The constraint equations describe the loop closures and are given by [8] \u03c6(q) = \u239b \u239d rp1 \u2212 rl1 rp2 \u2212 rl2 rp3 \u2212 rl3 \u239e \u23a0 . (44) Considering two of the base joints, q1, q3, as well as the elbow joints, q2, q4, and q6, to be actuated, we have k = 5 and Sp = \u239b \u239c\u239c \u239d 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 1 \u239e \u239f\u239f \u23a0 . (45) We will define the task to control the position of the platform (see Fig. 4(right)) while its orientation is uncontrolled. That is, m = 2, N = 1, and x (q7 q8 )T . (46) We will specify the reference value as xd = (\u22120.25 2.75 )T . (47) The null space torque will be specified to be zero (\u03c4N = 0). Additionally, we wish to specify the constraint forces at the interface of rp1 and rl1 . These correspond to \u03bbc ( \u03bb1 \u03bb2 )T . (48) We will specify the reference value as \u03bbcd = ( 25 sin(t/50) 150 cos(t/200) )T . (49) The remaining constraint forces will be unspecified. Thus, Sc = ( 1 0 0 0 0 0 0 1 0 0 0 0 ) (50) and Su = \u239b \u239c\u239c \u239d 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 \u239e \u239f\u239f \u23a0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001233_amr.874.57-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001233_amr.874.57-Figure3-1.png",
        "caption": "Fig. 3. Computation kinematic model of 3-axis hydraulic parallel manipulator.",
        "texts": [
            " The hydraulic cylinders was placed on the fixed base about circumradius R = 250 [mm], the rod cylinders are connected by the joints with the moving platform about circumradius r =130 [mm]. The maximum length of the hydraulic axis carries out Limax = Limin + h, where Limin = 425 [mm] is a initial length. The closed-loop kinematic chains of the hydraulic TPM create structure 3-RRPRR, in which revolute joints R and prismatic joints P step out. The 3D kinematic model and kinematic structure of the 3-RRPRR hydraulic TPM was shown in Figure 2. Kinematic of the hydraulic parallel manipulator The computational kinematic model of 3-axis hydraulic parallel manipulator shown in Fig.3. For the contour error of the 3-axis hydraulic parallel manipulator the 6-DoF kinematic model have been considered. It includes both the platform position variables xp, yp, zp and also rotational RPY angles , ,  . The RPY (Roll-Pitch-Yaw) angles defined with respect to three successive rotations about the fixed X, Y, Z axes [4]. The problem of inverse kinematics is to find the hydraulic cylinders elongation Li, given the position (xp, yp, zp) and orientation error (, , ) of the TCP. To solve the inverse kinematics problem vectors ri and Ri and matrix R A B are given, then vector Li from point A A i to point B A i can be calculated as: RrpL ii A Bi R  (1) where: p= \u2013 vector coordinates from point B in moving frame to point A in reference frame, ][ zyx ppp T p , R A B \u2013 rotation matrix from the moving frame to the reference frame for given values of RPY angles, ri \u2013 vector coordinates of B B i in the moving frame (B, X, Y, Z):  0 ii srcr T i r , r \u2013 radius of circumradius of the moving platform, Ri \u2013 vector coordinates of A A i in the reference frame (A, X0, Y0, Z0):  0 ii sRcR T i R , R \u2013 radius of circumradius of the fixed plate. As shown in Fig.3, all connection points Ai of the bottom fixed plate and connection points Ai of the top moving platform are find in the specific locations, as follows: for i = 1,2,3, then 1= 0 o , , . The length of vector Li can be obtained as: LL i T iiL  (2) The parallel mechanism has the characteristic of closed loop, so the constraints of the mechanism movement could be expressed as f(L,q)=0. Then the inverse kinematic solution of a parallel manipulator could be expressed as: qL dJd  (3) where: Jacobian matrix J is obtained as JJJ qL  1 , L is a vector which indicates the variation of hydraulic linear axes,  LLL T 321L , q is a vector which indicates the position and rotation of the moving platform,  zyx ppp T q "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001049_chicc.2014.6897031-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001049_chicc.2014.6897031-Figure1-1.png",
        "caption": "Fig. 1: Definition of the earth-fixed XEOYE and the bodyfixed XAY coordinate frames",
        "texts": [
            " Section 2 introduces some basics of ship mathematical model and the neural network. Then Section 3 designs the neural network observer, and the observer estimation errors and the observer weight estimation errors is proved to be uniformly ultimately bounded based on Lyapunov theory. In Section 4 an ex- ample is given to show the performance and effectiveness of the proposed neural network observer. Finally, conclusions are presented in Section 5. The reference coordinate frames of ship motion is illustrated in Fig.1, the earth-fixed reference frame is denoted as XEOYE and the body-fixed frame is XAY . The coordinate origin O of the earth-fixed reference frame is the original position of ships. The coordinate origin A of the body-fixed frame is located at the gravity center of the ship. Assume that the ship has an XZ plane of symmetry; surge is decoupled from sway and yaw; heave, pitch and roll modes are neglected. The nonlinear motion equations of a surface ship in the presence of disturbances can be expressed as [1,2] { \u03b7\u0307 = R(\u03c8)\u03bd M\u03bd\u0307 = \u2212D\u03bd + \u03c4 + w(t) (1) where the vector \u03b7 = [x, y, \u03c8]T represents the positions x, y and the yaw angle \u03c8 of the ship in the earth-fixed frame, \u03bd = [u, \u03c5, r] T represents the surge velocity u, sway velocity \u03c5 and yaw velocity r of the ship in the body-fixed frame, and \u03c4 = [\u03c4c1, \u03c4c2, \u03c4c3] T is the control input vector and moment provided by the propulsion system consisting of the surge force \u03c4c1, the sway force \u03c4c2 and the yaw moment \u03c4c3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000768_icuas.2014.6842236-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000768_icuas.2014.6842236-Figure1-1.png",
        "caption": "Figure 1. Hexacopter",
        "texts": [
            " However, in harsh environments failure situations can arise [4] and in these cases the control system of the UAV should be able to detect and counteract these failures preventing the aircraft from crashing. It is clear that the development of a self-learning and self-adaptive flight controller system is a primary step in the UAVs, in order to manage the aircraft when flight conditions change. A. The Proposed Approach In this paper a novel flight control technique, based on a Neural Network, for a hexacopter (depicted in Figure 1) will be shown. Moreover, a deep analysis will be carried out in order to identify the appropriate software/hardware architecture and the main components characterizing the flight controller. Through the results of this analysis, a development on a real embedded prototyping board will be presented. In order to check the validity of the proposed approach, the errors which affect the NN will be analysed and measured. In Section II main related works regarding to flight controllers of UAVs are shown"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure11-1.png",
        "caption": "Fig. 11 Two-carrier planetary gear train with two compound and four external shafts\u2014changegear train (gearbox)",
        "texts": [
            " The directions of the relative powers Prel I and Prel II in both component trains that are important for the determination of the efficiency are shown also. In the gear trains with internal power circuits it is shown that with one and the same structural scheme, the kinematic schemes may differ (Fig. 7). Figures 10 and 11 show two examples of the compound two-carrier planetary gear trains with two compound shafts and four external shafts. The brakes used in these gear trains can be located on different shafts\u2014single or compound. The Fig. 10 shows a reverse gear train, carrying forward and backward. The Fig. 11 shows a change-gear (gearbox), carrying out two gear-ratio steps (speeds). Figure 12 shows three cases of using two-carrier compound planetary gear trains with a one compound shaft, but with four external single (i.e. not-compound) shafts. Figure 13 shows the structural schemes of three-shaft three-carrier and fourcarrier compound gear trains. The first are used as reducers or multipliers, i.e. with F = 1 degree of freedom, while the second\u2014as change-gears. Figure 14 shows three-carrier compound gear train that works as a multiplier in the powerful wind turbine (Giger and Arnaudov 2011)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003610_ecce44975.2020.9235657-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003610_ecce44975.2020.9235657-Figure1-1.png",
        "caption": "Fig. 1. Proposed drive system integrated MMSG and multiple high speed motor for in-wheel motor system.",
        "texts": [
            " Since the high speed rotors have very small rotor diameter, it can drive in the high speed region because the Mises stress generating in high speed rotor can be decreased. Additionally, it realizes the high conversion efficiency even in high speed drive because the eddy current loss in the magnets and core loss can be much decreased owing to few harmonic fluxes in the flux transmission. Moreover, it can obtain the high torque density because the fundamental flux component of the high speed rotor contributes to generate the output torque and the transmitted torque can be increased in proportion to the number of multiple high speed rotor. Fig. 1 shows the proposed drive system. In the final target, the drive system of the MMSG and multiple high speed motors is applied to an EV\u2019s in-wheel 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 68 Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 18,2020 at 23:48:47 UTC from IEEE Xplore. Restrictions apply. motor system. The compact electromechanical drive system in which the motor, gear, inverter, and controller are integrated is realized by setting the inverter and controller to the space in the inner of high speed rotors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001628_ilt-04-2015-0049-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001628_ilt-04-2015-0049-Figure9-1.png",
        "caption": "Figure 9 Hypothetical labyrinth passages for the calculation of leakage rate in the radial direction",
        "texts": [
            " The real contact width is w1, and is supposed to be same with the circumferential contact width. For the calculation of the radial leakage rate, the leakage paths are simplified to be a similar labyrinth passage, as shown in Figure 8, and wr, hr and Lr are its width, height and length, respectively. wk is the breadth of the rough peaks, and it is regarded to be equal to Dm/nc. z(x) is the expression of the Abbott bearing surface curve shown in Figure 7. pi and po are the pressures at the inner and outer diameters of the metal gasket seal, respectively (Figure 9). wr is related with the embedded depth that the rough peaks embed the surface of the metal gasket seal, and can be expressed by: wr Dm 2nc 1 h1 hf (9) When h1 hf, hr varies periodically, and can be calculated by: Figure 8 Diagrammatic sketch of the interface of the surfaces of the metal gasket seal and the flange in the radial direction Leakage model of metallic static seals Chuanjun Liao, Xibao Xu, Hongrong Fang, Hongrui Wang and Man Man Industrial Lubrication and Tribology Volume 67 \u00b7 Number 6 \u00b7 2015 \u00b7 572\u2013581 D ow nl oa de d by U N IV E R SI T Y O F V IR G IN IA A t 1 7: 56 3 0 Ju ne 2 01 6 (P T ) hr hf h1 k 1)wp x wk/2 k 1)wp h0 h1 z k 1)wp x wk/2 k 1)wp x k 1/2 wp h0 h1 z kwp x k 1/2 wp x kwp wk/2 hf h1 kwp wk/2 x kwp wk Dm/nc k 1, 2 \u00b7\u00b7\u00b7 nr (10) When h1 hf, the rough peaks have been embedded into the softer coat of the metal gasket seal, the radial leakage paths have been closed, so the radial leakage rate can be considered to be zero"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003510_icarm49381.2020.9195311-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003510_icarm49381.2020.9195311-Figure4-1.png",
        "caption": "Fig. 4. The implementation of pheromones and robots in the simulated environment. Sensors of the agent are labelled with j \u2208 {0,1,2,3} corresponding to Y+,X+,Y\u2212,X\u2212 in this top view figure. The world\u2019s and robot\u2019s coordinates are labelled with X0,Y0 and X ,Y , respectively.",
        "texts": [
            " The successful deployment will be verified if all the agents have stopped and released the repellent pheromone at the desired positions before exceeding the defined experimental time Tout . The deployment strategy is implemented by Python and executed in the Jupyter iPython. The environment is composed of the following sections. 1) Arena Setup: We define a simulated rectangular arena with 0.8m width and 1.43m length (similar to the size of our real robot arena), within which the simulated pheromones are released and the simulated robots must begin exploring. An example of the pheromone implementation is shown by a heat map in the left part of Figure 4, where the darker colour represents a higher concentration of pheromone. 2) Simulated Robot: The robot is simulated as a 4cm diameter circle depicted in Figure 4. Four colour sensors are simulated as four points distributed in a 2\u00d72 array with 3cm diagonal distance to efficiently and conveniently get the instantaneous gradient of the pheromone. The concentration of the pheromone sensed by each sensor is calculated by (3)(2) and the average value is regarded as the current pheromone concentration sensed by the robot. The positive gradient direction of the pheromone in the robot\u2019s coordinates can be computed by (6), where i \u2208 {1,2} indicates two channels 978-1-7281-6479-3/20/$31"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002874_s12083-020-00878-6-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002874_s12083-020-00878-6-Figure2-1.png",
        "caption": "Fig. 2 Body-fixed frame and inertial frame of AUV",
        "texts": [
            " In Section 2, problem of AUVs are formulated, that is multi-input multi-output (MIMO) nonlinear systems. In Section 3, the analysis of polynomial regressors is described in detail, that extended the shortlength vectors to the polynomial regressors and improve control performance. In Section 4, the effectiveness of datadriven control is illustrated via simulation using the AUVs, that based on polynomial regressors. In general, the AUVs can be controlled advance, retreat, sideways, snorkel, and yaw with three propellers, which can rotate by 90 degree, as shown in Fig. 1. Figure 2 represents a few parameters of AUVs frame. The origin of inertial frame is O1, which is selected with arbitrary point on the sea level. The axis of O1X and O1Y are vertical on sea level. Furthermore, the O1Z axis is vertical to the sea level. The origin of body-fixed frame is O2, which is selected with the barycenter of AUVs. The advancing direction of the AUVs is represented as O2X0, O2Y0 and O2Z0 represent sideways direction and dive direction respectively. The AUVs angles of heeling, trim, and yaw are represented with \u03d5, \u03b8 and \u03c8 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003185_012049-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003185_012049-Figure8-1.png",
        "caption": "Figure 8. Flux distribution over healthy Induction Motor. (a) Flux lines. (b) Flux density.",
        "texts": [
            " Figure 7 delineates the spatial FFT range of airgap outspread motion thickness for the solid and flawed motor. Aside from the central part, different noises in solid motor range are because of openings, windings and inverter harmonics. The expansion in consonant substance because of static erraticism can be seen from Figure 7(b) contrasted with solid motor range. Figure 6(a) ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012049 IOP Publishing doi:10.1088/1757-899X/872/1/012049 The flux distribution of a fine fettle IM appears in Figure 8. It is seen from Figure 8(a), the dispersion of magnetic lines over the solid motor is balanced over the post pitch and each shaft is found at an ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012049 IOP Publishing doi:10.1088/1757-899X/872/1/012049 attractive pivot of 360\u02da/p geometrical degrees, where p is the number of posts. Whenever moment the bend of the periphery is \u03c0D/p for all the shafts, where D is that the inner distance across of the stator. Figure 9 shows the attractive field conveyance of IM under the static unusual condition with a seriousness level of 43%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003203_i2mtc43012.2020.9128407-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003203_i2mtc43012.2020.9128407-Figure1-1.png",
        "caption": "Fig. 1: test rig",
        "texts": [
            " Although each indicator will be affected by the failure of the blade individually, in order to improve the accuracy of the diagnosis, combining all the indicators is undoubtedly a better choice. One way to do this is using clustering algorithm. Since the variation of our indicators is relatively simple, in this paper we adopt a mature clustering method called Kmeans. In this section, the situation of our test bench will be briefly described and the procedure of the above method in crack diagnosis is illustrated. The test rig is shown in Fig.1, two air flow pipe are installed 180 degree apart which are not shown in the figure. 16 blades can be installed on this disk, but in order to reduce noise, only 8 blades were used with two of them have surface abrasion or crack at the bottom of the blade(Fig.2). The parameters of this test are shown in TABLE1. A structural analysis(Fig.3) of the normal and faulty blades are firstly demonstrated to observe the change of frequencies when faults exist. The frequency for normal blade is 365.07Hz, while for blade with crack and abrasion at the bottom are 350"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001113_10426914.2014.901531-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001113_10426914.2014.901531-Figure4-1.png",
        "caption": "FIGURE 4.\u2014Microhardness distribution on cross section.",
        "texts": [
            "98 14 21 485 16.0 Brass 934 0.071 8.5 108.9 390 20.6 FIGURE 2.\u2014Pulse waveform of laser welding. FIGURE 3.\u2014Penetration depth-to-weld width ratio. D ow nl oa de d by [ C hu la lo ng ko rn U ni ve rs ity ] at 1 8: 59 3 1 D ec em be r 20 14 where H is the penetration depth and B is the weld width, as shown in Fig. 3. It would form a narrow and deep welded joint when u has a higher value. The test line of the microhardness in the welded joint was along the dash line marked from brass to stainless steel in Fig. 4. The weld penetration depth and weld width with peak power and pulse duration under the condition of constant pulse energy were investigated. Experimental parameters are shown in Table 2 and the penetration depth and weld width value of Test No. 1 to No. 9 are shown in Fig. 5. Figure 5 shows the various penetration depths and weld widths in different parameters of test Nos. 1\u20139 (Table 2). As can be seen from Fig. 5, penetration depth value of Test No. 1 is less than Test No. 2 and Test No. 3. The reason may be that the peak power of Test No",
            " This was also an important reason of the crack formation. Moreover, as it was seen, cracks increases by increasing peak power in Fig. 9. The reason was that as the peak power increases, heating and cooling rates in the interface also increase dramatically. The increasing heating and cooling rates resulted in greater residual tensile stress and stress concentration, which bring out cracks in the welded joint easily. The test line of the microhardness in the welded joint was along the dash line marked from brass to stainless steel in Fig. 4. The results are the average value of testing for three times, as shown in Fig. 10. Microhardness in the weld of sample 2 and sample 9 are higher than that of sample 7. For sample 2 and sample 9, welded joint microhardness approximates base metal stainless steel because of the penetrated stainless steel in the welded joint. However, for sample 7, the microhardness decreased slightly in the welded joint because of the inadequately mixing of brass and stainless steel in welded joint. Figure 11 presents the respective shear failure force on each of the specimens"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure1.10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure1.10-1.png",
        "caption": "Fig. 1.10 Schematic of reciprocating linear generator system [24]",
        "texts": [
            " . . . . . . . . . . . . . . . . . 5 Figure 1.5 Linear machines for applications of rail transportation . . . . . 6 Figure 1.6 Winding arrangements for tubular PM linear machines [19] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 Figure 1.7 An axial flux permanent magnet motor [21] . . . . . . . . . . . . 9 Figure 1.8 Structure of the moving-coil-type linear DC motor [22] . . . . 10 Figure 1.9 Structure of the moving-coil-type linear DC motor [23] . . . . 10 Figure 1.10 Schematic of reciprocating linear generator system [24] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.11 A linear vibration-driven electromagnetic micro-power generator [26] . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.12 An improved axially magnetized tubular PM machine topology [28] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 Figure 1.13 Sectional view of a brushless DC linear motor [29] . . . . . . . 12 Figure 1",
            " presents a tubular linear brushless permanent magnet motor as shown in Fig. 1.9 [23]. It has a slotless stator to provide smooth translation without cogging. In this design, the magnets in the moving part are oriented in an NS-NS SN-SN fashion which leads to higher magnetic force near the like-pole region. Wang et al. describes the design and experimental characterisation of a reciprocating linear permanent magnet generator developed for on-board generation of electrical power fix telemetry vibration monitoring systems as shown in Fig. 1.10 [24, 25]. The two axially magnetised sintered NdFeB magnets and the mild-steel pole pieces give rise to an essentially radial magnetic field in the region occupied by the generator winding. The moving permanent magnet plunger is supported at each end by beryllium copper disc springs, which provide a high degree of radial stiffness, and an axial stiffness that can be accurately controlled by appropriate design of the spiral grooves. 10 1 Introduction Buren et al. presents the design and optimization of a linear electromagnetic generator suitable to supply power to body-worn sensor nodes [26]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001167_s00170-014-6109-8-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001167_s00170-014-6109-8-Figure11-1.png",
        "caption": "Fig. 11 Sets of isosurfaces for different constraint values (Ra\u22640.6 \u03bcm)",
        "texts": [],
        "surrounding_texts": [
            "4.1 Regression model of multiple constraints The cutting law could be defined by the relation between the cutting force and the cutting parameters. The cutting speed, teeth parameters of broaching tool (pitch, teeth number), and cutting edge radius remain the same in order to investigate the effects of tool-edge geometry parameters (RPT, rake angle, and relief angle) using the separation of variables. The different values of RPT and rake angle could change the uncut chip thickness directly. Rise per tooth and rake angle could affect the cutting force in the cutting zone. Relief angle has an important effect on the friction force in broaching process. Therefore, the broaching force per cutting edge also could be given as followed. F \u00bc f \u03b4; \u03b3;\u03b1\u00f0 \u00de: \u00f04\u00de Table 7 ANOVA table for the regression model F\u2019 Source Degree of freedom DOF Sum of squares SS Variance MS F test F0.05(9,5) Regression 9 267524 29725 18.90 4.77 Residual error 5 7862 1572 Total 14 275386 where \u03b4, \u03b3, and \u03b1 are RPT, rake angle, and relief angle, respectively. The regression equation of cutting force model could be proposed and obtained using the second-order polynomial model. In addition, the second-order regression equation of cutting force per cutting edge could be established using the CCF and RSM techniques as F \u00bc k0 \u00fe k11\u03b1 \u00fe k12\u03b3 \u00fe k13\u03b4 \u00fe k21\u03b1 2 \u00fe k22\u03b3 2 \u00fe k23\u03b4 2 \u00fe k31\u03b1\u22c5\u03b3 \u00fe k32\u03b1\u22c5\u03b4 \u00fe k33\u03b3\u22c5\u03b4 \u00f05\u00de where k0, {ki,j| i,j\u2208 [1, 2, 3]} are empirical constants. Therefore, the broaching force per unit length of one cutting edge could be expressed as F 0 \u00bc F=w: \u00f06\u00de where w is the width of cutting edge, which is also equal to a tooth\u2019s width of surface broaching tool. The response surface of broaching force per unite length is a kind of multi-parameter model which could be expressed as an equation. Meanwhile, in order to obtain the equation, the regression method on the basis of the experimental data was used for constructing regression models. The regression model of the resultant broaching force per unit length of one cutting edge is presented in Eq. 7. F 0 \u00bc \u22120:29\u03b32 \u2212 30:7\u03b12 \u00fe 83812\u03b42 \u2212 5:96\u03b3\u03b1 \u2212 53:3\u03b3\u03b4 \u2212 376\u03b1\u03b4 \u00fe 46:5\u03b3 \u00fe 422\u03b1 \u2212 7233\u03b4 \u2212 981 \u00f07\u00de The broaching force per unit length could be obtained experimentally and could produce similar values predicted from Eq. 7. In addition, it is needed to indicate the accuracy of the Eq. 7 by ANOVA analysis of the regression model. ANOVA indicates that broaching force model is adequate as P value of lack-of-fit is significant (P=0.002<0.05). As listed in Table 7, the calculated F statistic value (18.90) is greater than the standard table F0.05 (9, 5)=4.77. It is indicated that the regression fitting formula has a good agreement with the experimental curve. As depicted from ANOVA analysis in Table 7, it is indicated that the equations fit well with the experimental data. The regression curve and experimental data were also drawn in Fig. 5, which was with small error. The three cutting edge parameters (rake angle \u03b3, relief angle \u03b1, and RPT \u03b4) are subjected to the constraint of cutting force (i.e., the broaching force per unit length of one cutting edge, F\u2019), which are selected as independent design variables (X1, X2, and X3). The constraint F\u2019 with different values (F\u2019\u2264F\u20190) could to be represent by response surfaces. Furthermore, the broach tool geometry parameter [X 1, X 2, X 3] to satisfy a requirement of the constraint (F\u2019\u2264F\u20190) can be regarded as the points on response surfaces. For example, if the design constraint F\u2019\u2264500 N/mm, the parameter [X 1, X 2, X 3] could be selected from the series of isosurfaces, as depicted in Figs. 6 and 7. In order to meet the specified condition F\u2019\u2264F\u20190, X=[X1, X2, X3] can be selected from the number-setXset. Similarly, the response surfaces can also be obtained to satisfy the conditions of Ra\u2264Ra0 and S\u2264S0. Different cutting edge parameters of broach tool geomery have different effects on the broaching force. Figure 8 depicts main the effects of cutting edge parameters on average values F\u2019 obtained by ANOVA analysis. These figures show the effects of cutting edge\u2019s parameters on cutting force and roughness along with increase of rake angle, relief angle, and RPT. There is positive correlation among broaching force with RPT and relief angle. In contrast, there is a nonlinear correlation between broaching force and rake angle. Also shown in Fig. 8, it is further made clearly that the RPT is more effective than rake angle and relief angle. It is also important to note that the effect of relief angle is relatively smaller than rake angle. Furthermore, as seen in Fig. 8, it is observed that the interval range of rake angle from 15\u00b0 to 18\u00b0 will lead to direct decrease of resultant broaching force. The reason is that the large rake angle may play a critical role in reducing the broaching force. However, it is also indicated that when rake angles increase from the range of 18\u00b0~22\u00b0, broaching forces continually increase due to the sharp cutting edge tending toward damaged. Moreover, the small radiating area of sharp cutting edge will also cause wear. It can be concluded that larger tool rake angle may not always generate smaller broaching forces. When suffering shock and vibration, large rake angle will weaken tool edge strength instead of decrease Table 8 ANOVA table for the regression model Ra Source Degree of freedom DOF Sum of squares SS Variance MS F-Test F0.05 (9,5) Regression 9 0.363269 0.040363 32.22 4.77 Residual error 5 0.006264 0.001253 Total 14 0.369533 broaching force. In addition, it is clear from Fig. 8 that broaching forces change little with tool relief angle increasing from 3\u00b0 to 6\u00b0, which is less sensitive to the variation of tool relief angle. A positive factor affecting broaching force of relief angle is found over the entire evaluation domain from 3\u00b0 to 6\u00b0. Broaching forces increase gradually and the rate tends to be slow with the further increase of relief angle, especially when relief angle reaches 4\u00b0. Although the increase of relief angle is beneficial to reduce friction, it is also possible to employ the optimum effect of further minimizing broaching force. It appears that the condition of scattering heat in the cutting zone will also be worse correspondingly. Likewise, the broaching force and tool wear are also be affected by the coupling of cutting heat. It also concluded that the effects of relief angle on broaching are not great, but reasonable relief angle at 3\u00b0 and 4\u00b0 can reduce friction between tool and machined surface, resist shock and vibration as well as improving tool life. Generally, larger RPT obviously induces larger broaching force. It is illustrated in Fig. 8 that the values of RPT are correlated positively with broaching forces. Furtherly, as shown in Fig. 8, broaching forces change slowly with the increase of RPT in the range from 0.04 to 0.08 mm. When the values of RPT go beyond 0.08 mm, broaching forces increase markedly. Seen from the performances at different cutting edge parameters, the parameters chosen properly will benefit for reducing the broaching force. Base on the above analysis, the ideal combination of cutting edge parameters for lowest broaching cutting force should consist of moderate rake angle minimum RPT and relief angle. Similarly, the regression equation of broaching surface roughnessRa can also be obtained. And the quadratic response surface model depicting Ra can be also expressed as a function of design variables, such as \u03b4, \u03b3, and \u03b1 as given below in Eq. 8. Ra \u00bc 0:00293\u03b32 \u00fe 0:0086\u03b12 \u00fe 140\u03b42 \u2212 0:00172\u03b3\u03b1 \u2212 0:0582\u03b3\u03b4 \u2212 0:692\u03b1\u03b4 \u2212 0:120\u03b3 \u2212 0:028\u03b1 \u2212 16:2\u03b4 \u00fe 2:61 \u00f08\u00de where \u03b4, \u03b3, and \u03b1 are RPT, rake angle, and relief angle, respectively. Table 8 presents the results of analysis of variance (ANOVA) for Ra. It is evident that the calculated F value (32.22) is greater than the standard table F0.05 (9, 5)=4.77, which indicates the adequacy of the predictive model. The data of roughness obtained in experiment are depicted in Fig. 9. Correspondingly, the constraint Ra with different values (Ra\u2264Ra0) could also be represented by a mount of response surfaces. In addition, it is noted that the design parameter [X 1, X 2, X 3] of tool edge should also satisfy the constraint (Ra\u2264Ra0). Therefore, the design parameters can be selected from the coordinate values of points on the response surfaces, as given in Figs. 10 and 11. As depicted in Fig. 12, when rake angle under the range of 15\u00b0~18\u00b0, a increase of rake angle may lead to direct reduction of Ra. However, a further increase of rake angle especially in the interval range 18\u00b0~22\u00b0, Ra becomes increasing slowly. A similar phenomenon appears while changing the relief angle. It is demonstrated that sharp cutting edge is not always beneficial and helpful for the improvement of surface finish. The reason is that larger rake angle and relief angle deteriorate the strength of tool edge which easily affected by the shock and vibration during the broaching process. Compared with rake angle and relief angle, the changing of RPT has an influence on surface quality remarkably. Ra changes strikingly when RPT shifts from 0.04 to 0.12 mm. The increase in RPT makes the surface roughness poor when the RPT increases from 0.08 to 0.12 mm. Moreover, when RPT reaches 0.08 mm, a further rise may also result in a rapid increase of Ra. Additionally, as a consequence of ploughing, the value of roughness is large when the RPT is set as a relatively small value. Figure 13 indicates that the serious ploughing phenomenon occurs in the case of RPT setting to 0.04 or 0.05 mm, which will lead to poor surface roughness. In addition, enough tool edge strength of broaching tool is particularly important to improve tool life and machined surface quality. Furthermore, the structure strength of tool edge is another important constraint, which is also be determined by rake angle, relief angle, and RPT. The value of tool edge strength could be obtained by finite element methodology (FEM). In similar manner, the same CCF experiment design in the FEA simulation can also be applied to estimate the effects of RPT, rake angle, and relief angle on the strength of tool edge. As depicted in Fig. 14, the loads (Fx and Fy) measured by cutting experiments are applied on the nodes of cutting edge in FEAmodel. And the degree of freedom (DOF) of connected and fixed regions is set to zero (DOF=0). The regressionmodel of tool edge\u2019s structure strength S can also be obtained by RSM. The quadratic regression equation of maximum stress on the cutting edge is described as follows: S \u00bc \u22122:36\u03b32 \u2212 9:54\u03b12 \u00fe 22358\u03b42 \u2212 1:56\u03b3\u03b1 \u2212 11:0\u03b3\u03b4 \u2212 85:4\u03b1\u03b4 \u00fe 97:4\u03b3 \u00fe 123\u03b1 \u2212 2078\u03b4 \u2212 1053 \u00f09\u00de where \u03b4, \u03b3, and \u03b1 are RPT, rake angle, and relief angle, respectively. The adequacy of quadratic regression model is also estimated using F test. As depicted from ANOVA analysis in Table 9, it is also indicated that the equations fit well with the FEA-experimental data. The calculated F value (7.54) is greater than the standard table F0.05 (9, 5)=4.77, which indicates the adequacy of the predictive model. And the P value of lack-of-fit is also significant, which is smaller than the set probability 0.05 (P=0.019<0.05). In similar way, the design parameters of tool edge should also be subject to the constraint (S\u2264S0). As indicated in Figs. 15 and 16, response surfaces give vivid illustrations of the constraints. Response surfaces are formed by series of coordinate points of design parameters [X 1, X 2, X 3]. Therefore, the design parameters can be selected from the coordinate values of points on those response surfaces. 4.2 Selection of broach tool parameters based on CRSMMC The limitation of constraints\u2019 ranges is confirmed by practical application in industry production. Furthermore, these limit values should be determined by corresponding concrete application. The constraints (broaching force, surface roughness, and tool edge\u2019s structure strength) should satisfy the requirement of nominal power-load of machine tool, design standard of machined surface, and limitation of T15 steel\u2019s strengthen. Broach tool geometry parameters can be constrained by these certain ranges (F\u2019<F\u20190, Ra<Ra0, and S<S0). Therefore, broach tool geometry parameters are coupling constrained by many sets of response surfaces. Points on the response surface are considered as best choice of broach tool geometry parameters. Accordingly, the intersection of response surfaces, which satisfies the three constraints. Table 9 ANOVA table for the regression model S Source Degree of freedom DOF Sum of squares SS Variance MS F test F0.05(9,5) Regression 9 19897.6 2210.8 7.54 4.77 Residual error 5 1466.8 293.4 Total 14 21364.3 The limit values of these ranges are only supposed in order to explain how to apply CRSMMC. Furthermore, these limit values should be determined by corresponding concrete application. As shown in Figs. 7, 11, and 16, the constraints (F\u2019<F\u20190, Ra<Ra0, S<S0) are represented by several response surfaces in three dimensional spaces respectively. Correspondingly, when F\u2019, Ra, and S are exist simultaneously in a range (F\u2019<F\u20190, Ra<Ra0, and S<S0), a series of response surfaces were acquired as depicted in Fig. 17. Additionally, in this study, the three constraints in machining of heat-resistant steel X12CrMoWVNb N-10-1-1 were presumed Ra\u22640.6 \u03bcm, F\u2019\u2264500 N/mm, and S\u2264 2000 MPa. As indicated in the Fig. 17, the three constraints (Ra\u22640.6 \u03bcm, F\u2019\u2264500 N/mm, and S\u22642000 MPa) are be depicted by three given sets of response surfaces. It is noted that the intersection of the three sets of response surfaces can be obtained. As a result, the designing parameters [X 1, X 2, X 3] of broaching tool edge can be selected from the intersection of the three sets of response surfaces. As given in Fig. 17, the cutting edge parameters X=[X 1, X 2, X 3] can be selected from the intersection number- set XSet of response surfaces. Furthermore, the obtained intersection of response surfaces meeting variation ranges of three constraints is the point-sets of designing parameters [X 1, X 2, X 3]. Therefore, number-set XSet of the broach tool geometry parameters [X 1, X 2, X 3] constrained by three constraints are achieved in Fig. 18. In general, some points among all intersection points are selected to test and measure in experiment, in order to verify the validity of the method. However, it should be pointed out that the selected parameters from the intersection number-set XSet may be taken as the basis for the actual application. The determination of actual tool geometry parameters should be verified by experimental test. The six groups of tool edge parameters were chosen to conduct test verification. It is presented in Fig. 19a that partial points are chosen from all intersection points XSet given in Fig. 18. The six groups of tool edge parameters (i.e., from point 1 to point 6) are belong to the intersection number-set XSet. Measured values and errors are also given in Fig. 19b, c. As calculated in Fig. 19b and c, all the errors between the predicted values and measured values were controlled below 10.5 %. This demonstrated that the results acquired could be used as a reference for the decisionmaking of selection of tool edge parameters of broaching tool. In addition, it is also confirmed that CRSMMC is a creditable and effective method to provide valid evidences for acquiring optimum the geometry parameters of broaching tool. The flexible matching method CRSMMC of tool geometry parameters can also make tool\u2019s design more efficient."
        ]
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure2-1.png",
        "caption": "Fig. 2. Equivalent cumulated strains.",
        "texts": [
            " It is schematically shown in Fig. 1a. The initial ring is placed against the mandrel roll, which is forced toward the drive roll. The drive roll rotates continuously, reducing the wall thickness, and increasing the diameter. As shown in Fig. 1, the key challenge for this process simulation is that the material deformation occurs in a much reduced part of the computational domain. Preliminary calculations with Forge software [1] have provided non homogenous and abnormally high strain values \u00f0 e P 5\u00de (see Fig. 2), which appeared to be closely related to the contact treatment. Given the relative coarseness of the mesh (see Fig. 1b), the contact area is limited to very few nodes, so contact is not detected at the beginning of the process and temporal oscillations are observed. This is likely regarded to come from the coarse and non-smooth description of the cylindrical tools discretized by linear C0 triangular facets. In order to check this hypothesis, an analytical description of the tools is introduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001167_s00170-014-6109-8-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001167_s00170-014-6109-8-Figure17-1.png",
        "caption": "Fig. 17 Series of isosurfaces for three constraints",
        "texts": [
            "3 The limit values of these ranges are only supposed in order to explain how to apply CRSMMC. Furthermore, these limit values should be determined by corresponding concrete application. As shown in Figs. 7, 11, and 16, the constraints (F\u2019<F\u20190, Ra<Ra0, S<S0) are represented by several response surfaces in three dimensional spaces respectively. Correspondingly, when F\u2019, Ra, and S are exist simultaneously in a range (F\u2019<F\u20190, Ra<Ra0, and S<S0), a series of response surfaces were acquired as depicted in Fig. 17. Additionally, in this study, the three constraints in machining of heat-resistant steel X12CrMoWVNb N-10-1-1 were presumed Ra\u22640.6 \u03bcm, F\u2019\u2264500 N/mm, and S\u2264 2000 MPa. As indicated in the Fig. 17, the three constraints (Ra\u22640.6 \u03bcm, F\u2019\u2264500 N/mm, and S\u22642000 MPa) are be depicted by three given sets of response surfaces. It is noted that the intersection of the three sets of response surfaces can be obtained. As a result, the designing parameters [X 1, X 2, X 3] of broaching tool edge can be selected from the intersection of the three sets of response surfaces. As given in Fig. 17, the cutting edge parameters X=[X 1, X 2, X 3] can be selected from the intersection number- set XSet of response surfaces. Furthermore, the obtained intersection of response surfaces meeting variation ranges of three constraints is the point-sets of designing parameters [X 1, X 2, X 3]. Therefore, number-set XSet of the broach tool geometry parameters [X 1, X 2, X 3] constrained by three constraints are achieved in Fig. 18. In general, some points among all intersection points are selected to test and measure in experiment, in order to verify the validity of the method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000025_rnc.4734-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000025_rnc.4734-Figure1-1.png",
        "caption": "FIGURE 1 The relationship among the angles",
        "texts": [
            "aM = \u2212aM + VMq T\ud835\udefc \ud835\udefeM = \ud835\udf03 \u2212 \ud835\udefc, (1) where Fx = kF\ud835\udf0cV 2 Mcx (\ud835\udefc), Fz = kF\ud835\udf0cV 2 Mcz (\ud835\udefc,Mm), M = kM\ud835\udf0cV 2 Mcm (\ud835\udefc,Mm, \ud835\udeff), cz(\ud835\udefc, Mm) = cz1(\ud835\udefc) + cz2(\ud835\udefc)Mm, cm (\ud835\udefc,Mm, \ud835\udeff) = cm0 (\ud835\udefc,Mm) + c\ud835\udeffm\ud835\udeff, cm0(\ud835\udefc, Mm) = cm1(\ud835\udefc) + cm2(\ud835\udefc)Mm and T\ud835\udefc denotes the turning rate (incidence lag) time coefficient that is assumed as a constant. The definitions of variables or parameters can be found in Appendix A. Remark 1. For majority of interceptions, T\ud835\udefc is considered as a known slowly varying quantity and can be discussed as following equation: T\ud835\udefc = \ud835\udefc . \ud835\udefeM = V ( m \ud835\udf0cS )( \ud835\udf15cl \ud835\udf15\ud835\udefc )\u22121 , (2) where \ud835\udf0c = \ud835\udf0cV 2 M\u22152 denotes the kinetic pressure. In this paper, T\ud835\udefc is considered as a constant. Remark 2. The relationship among the angles \ud835\udf06, \ud835\udefeM, \ud835\udf03, and \ud835\udefc is shown as Figure 1. The angle of attack \ud835\udefc can be calculated via (1), and other angles can be measured by inertia sensors on the controlled missile. Remark 3. Assume that 4 fins are configurated as \u201c+\u201d structure, as shown in Figure 3, and then, fin deflection can be calculated by \ud835\udeff = (\ud835\udeff1 \u2212 \ud835\udeff3)/2. Because interceptor is an axisymmetric structure, fin deflection \ud835\udeff is equal to \ud835\udeff1 and \ud835\udeff3 numerically. Assuming that both the interceptor and the target are point masses, the planar homing engagement dynamics is depicted as in Figure 2, where M represents the interceptor and T represents the target, and other denotations can be found in the nomenclature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000129_ecce.2019.8911875-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000129_ecce.2019.8911875-Figure13-1.png",
        "caption": "Fig. 13. Proposal of rotor slit and design variables.",
        "texts": [
            " It is considered that the ratio of the end-leakage flux to the total flux linkage becomes large under the flux weakening control because only the flux in the core is reduced by the flux weakening. On the other hand, the result of the 2D FEA with Le is found to be in good agreement with the result of the 3D FEA and the experiment. From these results, the validity of the calculation method used in this paper is confirmed. IV. REDUCTION OF CROSS MAGNETIZATION To decrease the cross-magnetization effect and improve the characteristics of V-shape motor, a slit in the rotor core is proposed, as shown in Fig. 13. It is expected that the direction of the PM flux is changed by this slit and the cross magnetization effect decreases. An automatic shape optimization using FEA is carried out to determine the 6 dimensions x1, y1, x2, y2, W1, and W2 in Fig.13. Rosenbrock\u2019s method is applied for the optimization method because it often obtains better results with practical computational time [6]. The driving condition is set to be 500 r/min with maximum armature current under maximum torque control. The objective is to maximize the peak torque. The current angle  for the peak torque is determined by the iteration of 2D FEA. The mechanical stress caused by centrifugal force at high speeds is also estimated at each optimization step by coupling the stress FEA. The constraint condition is that the maximum Mises stress is less than 300 MPa at a over speed condition (8000 r/min) Fig. 14 shows the variation in the slit shape during the optimization. First, the slit inclines to the left side. Then, the width of the slit bottom at the right side is enlarged. Fig. 15 shows the variation in the torque with optimization iterations. The torque of the V-shape motor with the initial slit shown in Fig. 13 is considerably smaller than that of the V-shape motor without slit. On the other hand, it becomes larger by optimizing the slit as shown in Fig. 14. Fig. 16 shows the Mises stress distribution in the Vshape motor with optimized slit. It is confirmed that the maximum Mises stress is suppressed under the constraint condition. Fig. 17 shows the distributions of the total, PM, and armature reaction fluxes in the rotors with/without the optimized slit. It is observed that the optimized slit is nearly parallel to the total flux"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002050_s10999-016-9338-1-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002050_s10999-016-9338-1-Figure3-1.png",
        "caption": "Fig. 3 Contact state and elastic\u2013plastic region distribution of asperity with friction: a asperity side contact with friction and b distribution of elastic and plastic region (I elastic region, II plastic region)",
        "texts": [
            " Using Eqs. (8)\u2013(9), the elastic\u2013plastic normal contact stress in normal contact is given by r1 \u00bc 3F 2pr2 a 1 r2 p r2 a !1 2 \u00f015\u00de 3.2 Elastic\u2013plastic asperity contact model with friction When friction between two rough surfaces is included, the side contact occurs on the interface of two smooth asperities, as shown in Fig. 1b. Assuming that the contact region, whose distance is y2 from the contact point O3, is composed of the plastic (inner ellipse) and elastic region (outside ellipse annulus) (see Fig. 3). Similarly, y2 can be calculated from Eq. (11). The equivalent curvature radius of the two asper- ities at the contact point in XOY coordinate plane is R0 \u00bc R 1 \u00fe d2 0 R1 \u00fe R2\u00f0 \u00de2 \" #3 2 \u00f016\u00de The critical contact deformation of the asperity in the normal direction n0 is dn0c \u00bc pkH 2Et 2 R0 \u00bc pkH 2Et 2 R 1 \u00fe d2 0 R1 \u00fe R2\u00f0 \u00de2 \" #3 2 \u00f017\u00de When dn0 [ dn0c in the direction n0, the contact asperities are in plastic deformation stage. The radius of the contact area is r2 a0 \u00bc R0dn0 \u00f018\u00de Suppose that the plane O2O3Z1 is across the contact point O3 and is perpendicular to the plane XOY, the axis Z1 also crosses the contact point O3 and is perpendicular to the plane XOY",
            " It is shown in the isograms that the contact stress gradually decreases from the center to around. In Fig. 10a, the contact stress isogram in the plane at asperity normal contact state is a circle, which is similar to the elastic\u2013plastic region distribution of the asperity without friction by the present theoretical model in Fig. 2b. Likewise, the isogram at asperity side contact state is a ellipse in Fig. 10b, which is similar to the elastic\u2013plastic region distribution of asperity with friction by the theoretical model in Fig. 3b. The comparative results show that the normal contact stresses of the asperities on the tooth profiles are similar according to the present elastic\u2013plastic asperity contact model with or without the consideration of friction, or to the finite element models for asperity normal and side contact analysis. The results also indicate that the present theoretical model is well applied to elastic\u2013plastic asperity contact analysis on the rough tooth profiles of a spur gear drive. This study mainly focuses on an elastic\u2013plastic asperity contact model and its application for the micro-surface contact analysis of gear tooth profiles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002232_12.2177256-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002232_12.2177256-Figure2-1.png",
        "caption": "Figure 2. The above figures further illustrate the smoothing boundary layer concept [8].",
        "texts": [
            " Once within the existence boundary subspace, the estimated states are forced (by the SVSF gain) to switch back and forth along the true state trajectory. High-frequency switching caused by the SVSF gain is referred to as chattering, and in most cases, is undesirable for obtaining accurate estimates [16]. However, the effects of chattering may be minimized by the introduction of a smoothing boundary layer . The selection of the smoothing boundary layer width reflects the level of uncertainties in the filter and the disturbances. The effect of the smoothing boundary layer is shown in Fig. 2. When the smoothing boundary layer is defined larger than the existence subspace boundary, the estimated state trajectory is smoothed. However, when the smoothing term is too small, chattering remains due to the uncertainties being underestimated. Proc. of SPIE Vol. 9474 947409-4 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 09/30/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx Smoothers (whether fixed-interval, -point, or -lag) may be derived from the KF model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001210_ecce.2014.6953930-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001210_ecce.2014.6953930-Figure10-1.png",
        "caption": "Fig. 10. The test gears with (a) one-tooth-missing, (b) two-tooth-missing, and (c) gear crack.",
        "texts": [
            " 8 shows the experimental system setup, which consists of a 300-W PMSM driven by a variable-speed induction motor (IM) through two back-to-back connected gearboxes. These two gearboxes are two-stage gearboxes with 3 shafts and 4 gears. Their internal structure is shown in Fig. 9. The test gears are mounted at the input shaft of the test gearbox and pretreated by artificially generating various faults which were commonly observed in industrial systems, including one-tooth missing, two-tooth missing, and a gear crack (Fig. 10). One phase stator current of the PMSM is recorded with a sampling rate of 10 kHz for 210 seconds, during which the rotating speed of the PMSM is varied randomly in the range of 297 to 891 rpm. A total number of 1326 data records are collected, including 340 healthy cases, 318 one-toothmissing cases, 293 two-tooth-missing cases, and 375 gear crack cases. The two-stage drivetrain gearbox considered in this work includes 3 frequency components that are related to the shaft rotation and 2 frequency components that are related to gear meshing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000840_00423114.2015.1023319-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000840_00423114.2015.1023319-Figure4-1.png",
        "caption": "Figure 4. Finite element mesh of the centre plate arrangement unit (a) and contact surfaces (b).",
        "texts": [
            " The finite element models for the bolster and the body bolster beam were developed in analysis to determine the distributions of the contact stresses at the friction surfaces of the centre plate\u2013centre bowl pair. Some minor simplifications were made in the model, such as neglecting small elements, rounding, and facets that would not have serious influence on the distribution of contact stresses. The computer-aided design (CAD) model of the centre plate arrangement unit was developed using Kompas 3D CAD software. Then, the CAD model was imported to NEi Nastran v.8.3 finite element analysis (FEA) software for discretisation of the finite elements (Figure 4). DSMFEM software was used for further calculation of the contact pressures. DSMFEM is general-purpose FEA software used for the simulation of contact in engineering problems. Some applications are presented in [16,18\u201320]. The model includes four pairs of contact surfaces: two pairs of surfaces at the side bearings, the contact between the centre plate and the centre bowl, and the contact between the side surface of the centre plate and the bolster (Figure 4(b)). The contact surfaces of the centre plate were master surfaces, while the centre bowl\u2019s ones were slaves. The necessity of solving the contact problem implies some specific requirements for the finite element models. The first requirement comes from the idea of the force method: the nodes of the active and passive contact surfaces should strictly correspond to one another. Each pair of opposite nodes makes a contact pair. The active node (at the master surface from which the load is transferred to the slave surface) and the slave node should be able to coincide at the mid-surface as a result of deforming"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001876_cdc.2015.7402367-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001876_cdc.2015.7402367-Figure2-1.png",
        "caption": "Fig. 2. Model of a Two-Link Planar Arm",
        "texts": [
            " As expected from the physical limitations of the system, the obtained thresholds are positive if and only if ui \u2264 Gi(v) \u2264 ui \u2200v. In other terms, to ensure feasibility it is required to ensure that, at least in the currently applied reference v, the actuator torques must be able to compensate the gravitational terms. Having obtained the energy threshold associated to each constraint, constraint enforcement can be achieved by manipulating the applied reference according to equations (22) and (24)-(26). For the sake of providing some numerical simulations, consider the two-link planar arm illustrated in Figure 2, where the matrices in equation (28) are specified in (27). The robot is controlled using the PD with desired gravity compensation (29) with KP = [ kp1 0 0 kp2 ] + [ \u03b71 + \u03b72 \u03b72 \u03b72 \u03b72 ] KD = [ kd1 0 0 kd2 ] , (33) M (q) = [ \u00b51 + 2\u03bdc2 \u00b52 + \u03bdc2 \u00b52 + \u03bdc2 \u00b52 ] C (q, q\u0307) = [ \u2212\u03bds2q\u03072 \u03bds2 (q\u03071 + q\u03072) \u03bds2q\u03071 0 ] G (q) = [ \u03b71c1 + \u03b72c12 \u03b72c12 ] (27) \u00b51 = m1l 2 1+m2l 2 2 3 +m2l 2 1 \u00b52 = m2l 2 2 3 \u03bd = m2l 2 2 4 \u03b71 = ( m1 2 +m2 ) gl1 \u03b72 = m2 2 gl2. and kd1, kd2, kp1, kp2 > 0. Note that KP satisfies (30). The resulting Lyapunov function (31) is V (v, q, q\u0307) = 1 2 (q\u2212v)T [ kp1 0 0 kp2 ] (q\u2212v) + q\u0307TM (q) q\u0307 +\u03b71 S (v1, q1) + \u03b72 S (v1+v2 , q1+q2) where S (a, b) = sin (b)\u2212 sin (a)\u2212 cos (a) (b\u2212 a) + 1 2 (b\u2212 a)2 is a positive semi-definite function"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001125_s12206-013-1181-9-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001125_s12206-013-1181-9-Figure5-1.png",
        "caption": "Fig. 5. Single type pilot model for mooring winch system.",
        "texts": [
            " This means that the cable tension affects the vessel motions as an unpredictable input force. Then, the input force made from cable tension variation can be considered as a disturbance. Therefore, if we have the frequency and amplitude information about the uncertain cable motions, we can provide a useful control strategy to cope with the uncertainty. Based on this fact, the authors found the information required for designing the control system by experiment. To evaluate the proposed control strategy, a pilot model is made to simulate a mooring winch system as shown in Fig. 5. In Fig. 5, \u2462~\u2464 and \u2466 represent the vessel part which slides on the frame. The vessel part is connected to \u2460 (damper) and \u2461 (spring) which are equipped to cope with the passive cable motions as shown in Figs. 2 and 3. As explained earlier, the cable motions are identified from the experiments of winding and unwinding the winch. These actions induce various cable motions which give input force changes to the vessel, and thus are considered as uncertainties. First, let us define the parameter values of the vessel part given in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-20-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-20-1.png",
        "caption": "Figure 6-20. References for obtaining the allocation matrix of a non-coaxial bicopter",
        "texts": [
            " Notice that the control analysis of the center of the vehicle was only made with half of the motors. This is only possible while the copy-motors (in this case 1 and 3) are not placed far from the action-motors (in this case 2 and 4). Notice that the action part of the propulsion or allocation matrix is similar to a differential wheeled robot. Without the copy-motors of the coaxial system, if you increase the speed of motor 2 and reduce the speed of motor 4, in addition to torque on the X axis of the drone, you would also have torque on its Z axis. See Figure\u00a06-20. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 264 Therefore you would have a different system (remember the right- hand rotation criteria for the Z axis of the drone): Fz x z t t w w \u00e9 \u00eb \u00ea \u00ea \u00ea \u00f9 \u00fb \u00fa \u00fa \u00fa = - - \u00e9 \u00eb \u00ea \u00ea \u00ea \u00f9 \u00fb \u00fa \u00fa \u00fa \u00e9 \u00eb \u00ea \u00f9 \u00fb \u00fa 1 1 1 1 1 1 2 4 This implies a pseudoinverse problem and therefore a more complex computational optimization problem. Notice that, unlike a hexacopter which has more engines and only four possible independent movements, in this case you have fewer motors than feasible movements (only two of them to be used with three possible independent movements), falling into an underactuated robotics problem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002473_j.measurement.2016.11.034-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002473_j.measurement.2016.11.034-Figure2-1.png",
        "caption": "Fig. 2. Illustration of the first eigenmodes of the existing exp",
        "texts": [
            " Study of the existing uniaxial dynamometer The dynamic behavior of the existing device was first investigated by FEM modeling using the commercial software Abaqus n (a) and the tool-sample measurement configuration in (b). to calculate its rigidity and its eigenfrequencies. Due to the high inertia of the whole ballistic bench and thus to the important energy needed to excite it, the influence on the dynamic behavior of the whole structure was neglected. Only the interaction part of the experimental bench was modeled (Fig. 2). Faces A and B were fixed in the model (Fig. 2a). Elastic behavior was considered for all materials whose properties are reported in Table 1. For the force sensor only the mass and the stiffness were given by the manufacturer, therefore its material properties have been determined numerically. A model of the sensor, considering an equivalent homogeneous material, subjected to a compression load and clamped on one of its faces was used (i.e. the face C in Fig. 2b). The Young modulus E and the volumetric mass density q were then set in order to obtain the same stiffness and mass that those of the real sensor. erimental device. Nodes of surfaces A and B were fixed. Table 2 Resonance frequencies (a) and stiffnesses (b) obtained from experiments and numerical modeling. Mode Resonance frequency (Hz) Deviation (%) Numerical Experimental (a) 1 1687 1632 2.97 2 2132 1998 6.20 3 2545 2456 3.50 4 4117 4104 0.32 5 5831 5981 2.57 6 6488 6500 0.18 7 7025 6881 2.05 Stiffness (N/lm) (b) Cutting direction 13 Tangential direction 12 Normal direction 20 (a) (b) (d) (g) (e) Fig",
            " The newly op shown in (g). Table 2a presents a comparison between the system response resonance observed experimentally [16] and the eigenfrequencies obtained from the numerical modal analysis. From the FEM analysis the deformation is mainly in the cutting direction. The disagreement on the resonance frequency determination between these two approaches was above 6%. In fact, as shown by the first eigenmodes for the tangential and normal directions, the tool fastening was found to be too weak (as shown by Fig. 2c and d). This was essentially due to the cantilever mounting design and to the lack of rigidity of the preloading screw (diameter 8 mm). In addition, this model also shown a lack of stiffness of the device mounting in the tangential and normal directions (Fig. 1e and f). As shown by previous works [12,17] in high-speed contact measurements between an aluminium-based abradable material specimen and a 42CrMo4 steel tool, the cutting force could reach about 1800 N. The coating wear variation as a function of the specimen position in the cutting direction was measured to be about 69 lm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002866_tii.2020.2978771-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002866_tii.2020.2978771-Figure3-1.png",
        "caption": "Fig. 3. FDTD model of the experimental apparatus (only a selection of rebar positions are shown; the tank and some details of the antenna are omitted for illustrative purposes). N.B. measurements were made with a single rebar at each location in turn.",
        "texts": [
            " In common with many other GPR simulations [20]\u2013[23], a Gaussian shaped pulse was assumed with a center frequency of 1.5 GHz. A simple Gaussian shape is a good approximation, but may not be an entirely realistic representation of the real pulse, which is often generated by an avalanche transistor. A feed model consisting of a voltage source with internal resistance inserted in a one-cell gap between the two arms of the transmitter bowtie (the drive-point) was used. Fig. 2 shows the detailed FDTD mesh of the geometry of the antenna, and Fig. 3 shows the FDTD mesh of the experimental apparatus. A spatial discretization of \u0394x = \u0394y = \u0394z = 1 mm was chosen as a good compromise between accuracy and computational requirements. gprMax computes the spatial and temporal derivatives using a standard second-order scheme and this choice of spatial discretization also ensured that any numerical dispersion was adequately controlled. The Courant Friedrichs Lewy (CFL) condition was enforced which resulted in a time-step of \u0394t = 1.926 ps. The three emulsions and the tap water used in the experiments have frequency-dependent conductivities [13] which were modeled by fitting a Debye formulation [24]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000287_humanoids43949.2019.9035021-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000287_humanoids43949.2019.9035021-Figure3-1.png",
        "caption": "Fig. 3. Skeleton of the GOOGOL GRB3016 robot with coordinate frames in the home position.",
        "texts": [
            " In the update step, the state covariance matrix is updated by optimal Kalman gain K as follows: Kk+1 =Pk+1|kJ T k+1S \u22121 k+1 x\u0302k+1|k+1 = x\u0302k+1|k +Kk+1y\u0303k+1 Pk+1|k+1 =(I \u2212Kk+1Jk+1)Pk+1|k (26) where I is the identity matrix. Once the updating procedure is completed, the norm values of the state vector are calculated for every iteration. The EKF is reduced to the Newton\u2013Raphson method when Q and R are set to zero. A GOOGOL GRB3016 robot was used in this experiment to verify the proposed method. The robot can self-calibrate online in its working status. The nominal robot link parameters are shown in Table I. Fig. 3 shows the skeleton of the GOOGOL GRB3016 robot with all its coordinate frames and geometric features. The matrices R and Q for the KF can be determined by the adaptive method described in [28] and [29]. The matrices R and Q for the EKF can be calculated through the method in [30]. The GOOGOL GRB3016 robot with six DOFs requires 24 geometric parameters to be modeled. Each 3-D robot pose provides six model equations, as indicated by (19). Therefore, a unique computation of the 24 parameters requires a minimum of four pose measurements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000380_iros.2011.6095187-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000380_iros.2011.6095187-Figure1-1.png",
        "caption": "Fig. 1. Desired trajectories for window cleaning",
        "texts": [
            " This study aims to develop a high efficiency and high reliable glass cleaning robot. Thus, we adopt the model-based method for the control of the cleaning robot. For the efficient and reliable cleaning, the robots should follow accurately a desired trajectory which covers the whole window with less overlapping of the sweep motion. Considering the fact that most windows are rectangle one, trajectories with (a) horizontal parallel motion, (b) vertical parallel motion and (c) contour parallel motion as shown in Fig.1 are efficient ones. However, because water is used for the window cleaning, dirty water goes down and makes a cleaned glass dirty with vertical parallel motion trajectory and contour parallel motion one. Therefore, the robot follows Y. Katsuki, T. Ikeda and M. Yamamoto are with the Faculty of Engineering, Kyushu University, 744 Motooka, Nishi-ku, Fukuoka 819-0395, Japan katsuki@ctrl.mech.kyushu-u.ac.jp the horizontal parallel motion trajectory as the desired one in this study. On the trajectory control problem for the glass window cleaning robot, an attitude control method for a wallclimbing robot has been proposed [2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003642_amc44022.2020.9244337-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003642_amc44022.2020.9244337-Figure2-1.png",
        "caption": "Fig. 2. Planar manipulator modeled in simulations.",
        "texts": [
            " (32) In (32), gf stands for the cut-off frequency of the first-order low-pass filter Qf . When \u02c6\u0307\u03c3dis is obtained, x\u0308des may be selected as x\u0308des = \u02c6\u0307\u03c3dis \u2212\u03a8(\u03c3f ) (33) where \u03c3\u0307f + \u03a8(\u03c3f ) = 0 (34) defines the desired first-order dynamics of \u03c3f . Possible solutions are \u03a8(\u03c3f ) = K\u03c3f (35) \u03a8(\u03c3f ) = Ksign ( \u03c3f ) (36) where K is a diagonal matrix with positive diagonal entries. The proposed control method was validated in simulations. The robot was selected to be planar manipulator, with two actuated joints. The modeled manipulator is illustrated in Fig. 2. The kinematic and dynamic model of the manipulator are taken from [8]. The reader is advised to consider this reference for the details of the kinematic and dynamic model. The dynamic model from [8] is extended with \u03c4 ext, which is taken as the vector of external forces (actually torques in this case) acting on the joints due to interaction with an obstacle. The final model is having the form (1). The following notation is used for the manipulator\u2019s parameters. For i = 1, 2, mi is the mass of link i; Ii symbolizes the moment of inertia of the link i about an axis coming out of the page, passing through the center of mass of the link i; li stands for the length of link i; and lci represents the distance from the previous joint to the center of mass of link i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000826_oceans-bergen.2013.6608110-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000826_oceans-bergen.2013.6608110-Figure5-1.png",
        "caption": "Fig. 5. Definition of the \u03b8 angle.",
        "texts": [
            " Assuming that from the current point to the next way point the vehicle will move along a straight path with the fixed heading \u03b8, the next way point is calculated as minimum of the Renyi\u2019s entropy, among all the possible choice of heading angle for the straight paths S\u03b8, with length not exceeding Rmax and still belonging to sub-region Qi. Where, \u03b8 is the angle, in a clockwise direction, between x-axis of the navigation system reference and the straight-path from the vehicle current position to way-point at max length Rmax (see Fig.5). Formally: min\u03b8 HR(p, S\u03b8) = \u2212 ln \u222b S\u03b8 p2(x)dx S\u03b8 \u2282 Qi \u2200\u03b8 l(S\u03b8) \u2264 Rmax (18) where, S\u03b8 is the straight path in function of \u03b8, l(\u00b7) is the norm of the vehicle\u2019s straight path and Rmax is the maximum path length which one vehicle can execute from current position to way-point. Once the next way-point has been established, navigation to the way-point is determined using behaviours implemented through potential functions, as discussed in detail in [10]. Two potential functions have been defined: attraction and obstacle avoidance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000391_978-3-319-02294-9-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000391_978-3-319-02294-9-Figure11-1.png",
        "caption": "Fig. 11 The distribution of magnetic field in the air gap as provided by ferrite magnets",
        "texts": [
            " 7 Total energy consumption for particular machine tool components production The grafs show that the following parts of machine tool consume the most of total energy: sled with revolving stand, sled and stand. The results of the third LCA phase are shown in the form of partial graphs presenting environmental burden by particular machine tool components according to the chosen impact categories [15]. 70 M. Iskandirova et al. Fig. 8 Partial results of indicators for ozone layer depletion CO2 eq. Fig. 9 Partial results of indicators for ozone layer depletion CFC-11 eq. Fig. 10 Partial results of indicators for acidification SO2 eq. Fig. 11 Partial results of indicators for photooxidants creation C2H4 eq. Fig. 12 Partial results of indicators for eutrophication PO4 3- e Fig. 13 Legend to Figures 5 to 13 According to the results of this phase, it is possible to make following conclusion about components the most burdening the environment in chosen impact categories: sled with revolving stand, sled and stand the most contribute to global warming and eutrophication; headstock the most contributes to ozone layer depletion; sled with revolving stand, sled, stand and headstock the most contribute to acidification; platform the most contributes to photooxidants creation",
            "7 0 5 10 15 20 25 30 35 m a g n itu d e o f h a rm o n ic c o m p o n e n t [ T ] order of harmonic component [-] Spectrum of Winding's Magnetic Flux Density Fig. 10 The FFA of the ferrite solution The Comparison of the Permanent Magnet Position in Synchronous Machine 289 The maximum value of air gap magnetic flux density reaches values higher than 1T. However, these values could be found only in areas above the iron parts of rotor between magnets. The value of the fundamental harmonic component of magnetic flux density in the air gap is, according to FFA, slightly over 0.6 T only. The distribution of magnetic field is shown in the fig.11. As seen from fig. 9, the magnetic flux density distribution along the air gap forms a row with triangular shape, which is not suitable for use in synchronous machines. This magnetic flux density dependence would cause unsuitable harmonic components in induced-back electromotive force resulting in problematic control of the machine. [1] [5] [11] The solution of the permanent magnet synchronous motor, presented in this paper, shows that for machines with higher amount of poles with relatively small pole pitch are better to use inner storage of permanent magnets",
            " The aim was to find the maximum torque value for a drive used for the movement. The maximum torque value (Figure 10) was Mk = 1602,042 N.mm (16,02 kg.cm) in time 0,105 s. The obtained dependence of the torque has been used for the drive selection. The large torque value is caused by the starten of the drive from zero to maximum velocity. Fig. 9 Drive of the wheels Fig. 10 Movement simulation of the drive of wheels 548 M. Dovica et al. The chosen drive has been used for the simulation of other parts via FEM (finite element method) (Figure 11). Upper position of undercarriage has to be fixed, because the mechanism is not self locking. The weight of the robot may lead to the fall of the lower part of the undercarriage. For this purpose, the locking mechanism was designed. (Figure 12) This mechanism consists of servomechanism that moves the pin into the locking position. This solution is also suitable from the energy point of view. Final design of the robot is shown in the Figure 13. The overall dimensions are 280 x 351 x 175 mm and its weight is 3,3 kg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002866_tii.2020.2978771-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002866_tii.2020.2978771-Figure2-1.png",
        "caption": "Fig. 2. FDTD mesh of antenna model (main features annotated).",
        "texts": [
            " The excitation of the antenna\u2014pulse shape, frequency content, and feed method\u2014is important for the performance of the real antenna, and hence critical to capture in the model. In common with many other GPR simulations [20]\u2013[23], a Gaussian shaped pulse was assumed with a center frequency of 1.5 GHz. A simple Gaussian shape is a good approximation, but may not be an entirely realistic representation of the real pulse, which is often generated by an avalanche transistor. A feed model consisting of a voltage source with internal resistance inserted in a one-cell gap between the two arms of the transmitter bowtie (the drive-point) was used. Fig. 2 shows the detailed FDTD mesh of the geometry of the antenna, and Fig. 3 shows the FDTD mesh of the experimental apparatus. A spatial discretization of \u0394x = \u0394y = \u0394z = 1 mm was chosen as a good compromise between accuracy and computational requirements. gprMax computes the spatial and temporal derivatives using a standard second-order scheme and this choice of spatial discretization also ensured that any numerical dispersion was adequately controlled. The Courant Friedrichs Lewy (CFL) condition was enforced which resulted in a time-step of \u0394t = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000459_b978-0-12-415995-2.00013-1-Figure13.12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000459_b978-0-12-415995-2.00013-1-Figure13.12-1.png",
        "caption": "FIGURE 13.12 Simplified sketch of a basic principle in autotomy, i.e., self-amputation, seen in some lizards. When certain muscles contract, one of the vertebrae breaks along a predefined fracture line. When the predator pulls the tail, the tissue is torn along weak planes. The drawing is based on the description of autotomy in Ref. 44.",
        "texts": [
            " However, common to all groups is that autotomy occurs along predefined breakage lines, such as between specific leg joints in spiders or at the tail root in lizards [43]. The predefined breaking plane is particularly weak and gives a clean break with minimal external force, which ensures that the amputation is followed by a minimal loss of blood and in many cases also facilitates the ensuing regeneration. Sometimes, autotomy in lizards does not happen at the joints between two vertebrae but by bone fracture [44]. The core principles are illustrated in Figure 13.12 based on the description in Ref. 44. A fracture along a breaking plane in one of the vertebrae is activated by contraction of tail muscles. An external pull of the tail causes the remaining tissue to break along weak planes. The principles of autotomy could be used in many places\u2014for example, in fire protection of buildings and in flow control of sewer systems to avoid massive central overflows during heavy rains. Disassembly of industrial products such as cars and washing machines can also benefit from these principles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000003_s00366-019-00870-6-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000003_s00366-019-00870-6-Figure1-1.png",
        "caption": "Fig. 1 Schematic of stepped silo with four different wall thicknesses",
        "texts": [
            " For reasons of saving the potential cutting costs, the heights of most silo segments should be an integral multiple of the corresponding plate width value as far as possible, while the overall height of the silo should be kept still. To realize this goal, numerical processing techniques, such as generating a random number from a uniformly distributed set of discrete positive integers, linear normalization and linear interpolation, etc., were applied in this study. A simply supported flat-bottomed cylindrical silo, composed of rolled plates of four different wall-thicknesses (Fig.\u00a01), was modeled and optimized in this work. By reference to [8, 9], etc., the silo roof and hopper were neglected in the subsequent FEM model. The silo size was: height H = 22.6\u00a0m, internal diameter D = 3.2\u00a0m (H/D = 7.0625). The silo consisted of annealed aluminum alloy type 3003 plates (AA3003), which was assumed elastic\u2013perfectly plastic (without hardening) with the following properties: volumetric weight \u03b3 = 26.75 kN/m3, elastic modulus E = 69 GPa, Poisson\u05f3s ratio \u03bd = 0.3 and yield stress \u03c3y = 40\u00a0MPa",
            " At the silo bottom, the radial, circumferential, and meridional displacements and rotations were prevented, while at the silo top, the radial and circumferential displacements were fixed only. The 4-node thin shell elements with a reduced integration point and hourglass control (S4R) were employed. The finite element size was 10 \u00d7 10\u00a0cm2 and the total amount of finite elements was about 22,900. To assign the discrete wall-thicknesses, the silo shell geometry was firstly partitioned by three datum planes, which were positioned along the silo height direction by means of the parameters h1, h2 and h3 (labeled in Fig.\u00a01), respectively. After partition, according to the thickness values t1, t2, t3 and t4 (labeled in Fig.\u00a01), the wall-thickness of each silo segment was defined through the homogeneous shell section editor of ABAQUS. According to EN 1991-4:2006, the load on silo vertical walls is composed of a fixed load, called the symmetrical load (including the horizontal wall pressures and the wall frictional traction), and a free load, called the patch load, which shall be taken to act simultaneously. For the sake of safety and reliability, the loads of discharge process were used for silo nonlinear implicit dynamic FEA due to its abrupt dynamic characteristic",
            " The discrete characteristics of the size specification of the commercially available rolled plates set an unavoidable constraint for silo size optimization, which implies that there are only a few specific values of ti could be selected. In addition, the values of plate width are also discrete and a width value is often associated with a number of thickness values (reference to Table\u00a01). For reasons of saving the cutting costs, hi (i = 1, \u2026, N\u22121) should be an integral multiple of the corresponding width value as far as possible, while the overall height of the silo should be kept still. As depicted in Fig.\u00a01, the value of N was taken as 4 in this work. Particle swarm optimization (PSO) is a population-based stochastic and metaheuristic optimization algorithm inspired by social behavior of bird flocking or fish schooling and developed by Eberhart and Kennedy in 1995 [37\u201339]. PSO involves a swarm of particles that represent the potential solutions to the problem. A swarm consists of Np particles moving around an N-dimensional search space, each particle representing a potential solution. The ith particle is characterized by its position vector Xi= [xi1, xi2, \u2026, xiN] and velocity vector Vi= [vi1, vi2, \u2026, viN]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000267_icsmartgrid48354.2019.8990687-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000267_icsmartgrid48354.2019.8990687-Figure6-1.png",
        "caption": "Fig. 6. Measurement of electrically conductive part. (a) Measurement conditions. (b) Clamp fastening plates.",
        "texts": [
            " Hereinafter, the zone measuring an electric potential distribution is referred as the \u201czone measuring potentials.\u201d The area of the zone is 110 mm by 35 mm. The lid measuring potentials overlaps the four flanges of the metal box shown in Fig. 4. One flange is metal and the other flanges are insulated. The width, length, and thickness of the metal flange are 100, 30, and 1.2 mm, respectively. As shown in Fig. 5, the lid measuring potentials overlaps the metal box. When a voltage is applied to the lid and the box as shown in Fig. 5, a current can flow only through the metal flange. Fig. 6(a) shows measurement conditions. In case A, a 1 kg weight is on the lid. In case B, the lid and the metal flange are fastened by two clamps shown in Fig. 6(b). Fig. 7 is a schematic illustration of the electric potential distribution measurement method. The terminals of the lead wires attached to the grids are connected to relays, which are controlled by a personal computer (PC), and the connection between an amplifier and the grid is changed by the relays. The electric potential between a reference for potentials and the grid is amplified and input to an analogue digital converter in the PC. Based on the input data, the PC displays the potential distribution of the grids. An arbitrary grid is assigned as the reference for potentials. A dc current (10 A) is applied to the lid and the metal flange as shown in Fig. 6(a). Figs. 8 and 9 show the observed electric potential distribution for the cases A and B. On the basis of these distributions, current density distributions are calculated by (9). Figs. 10 and 11 show the distribution of the extracted current density Ie/(\u0394x\u0394y) (\u03c3 of (9) is the electric conductivity of steel [104 \u03a9\u00b7mm)] and electrically conductive parts. The x and y coordinates of Figs. 8\u201311 are shown in Fig. 3. The distributions of the extracted current density show that current flows through a comparatively small area (about 10 \u00d7 10 mm2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002341_ijrapidm.2016.078746-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002341_ijrapidm.2016.078746-Figure14-1.png",
        "caption": "Figure 14 Position of measure angles in ramp plate for angular deviation",
        "texts": [
            " The result suggests that it is better to print holes along z-axis than along x and y axes for improved results. Subsequently, the orientation of the part on AM machine becomes an important factor. Additionally, the graph clearly shows that average circularity deviation does not exceed 0.2 mm, which is in accordance with the earlier results (see Section 3.2.1). The ramp plate was the next part printed to investigate the angular deviation for wedge feature fabrication. The angular tolerance is defined as the maximum allowable deviation from a specified angle. Figure 14 shows the location of angles that were measured. Nine parts were printed to get a better repeatability for measurement of angles. The part angles were measured using a computer-controlled touch probe CMM. Figure 15 shows the uniformity in the sign of angles of all the printed parts. Owing to less amount of data, it is difficult to provide reasons for this trend. However, the graph shows that the angular tolerance is within 0.65\u00b0 for all the three angles. The bevel gear model possesses high level of rotational symmetry with smallest of differences in surface curvature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002254_s0081543816060171-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002254_s0081543816060171-Figure5-1.png",
        "caption": "Fig. 5. Gyration ellipsoid and the location of the center of mass for the Hess case.",
        "texts": [
            " We recall the geometric meaning of the corresponding restrictions on the parameters, assuming that all principal moments of inertia are different: \u2022 make a transformation from the angular velocities to the angular momenta: M = I\u03c9; \u2022 in the three-dimensional space of angular momenta, consider the level surface of the kinetic energy, the gyration ellipsoid (M, I\u22121M) = const; \u2022 draw a pair of circular sections passing through the middle axis of the gyration ellipsoid. In the Hess case the center of mass lies on the perpendicular to the circular section of the gyration ellipsoid (see Fig. 5). Now let us rewrite the constraint equation (2.1) as (I\u03c9,a) = 0, a = I\u22121e. It can be shown that if Ox1 is the middle axis of inertia, then the condition for the vector n to be perpendicular to the circular section of the corresponding gyration ellipsoid coincides exactly with the conditions for the existence of an integral with U = 0 and k = 1: I13 = 0, I11I22 = I222 + I223. (2.16) PROCEEDINGS OF THE STEKLOV INSTITUTE OF MATHEMATICS Vol. 294 2016 In addition, we now assume that the center of mass of the body is also located on the normal to the same circular section in the chosen coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001589_saci.2013.6608956-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001589_saci.2013.6608956-Figure5-1.png",
        "caption": "Figure 5. Correlation between dvqv frame, dq frame and stator frame",
        "texts": [
            " At ramp startup, under no load conditions, the voltage amplitude command V1 * is determined from (1, 2): V1 * = \u03bbPM \u00b7|\u03c9r *|, (11) C.-E. Coman et al. \u2022 V/f Control Strategy with Constant Power Factor for SPMSM Drives, with Experiments \u2013 148 \u2013 where \u03c9r * is the electrical speed reference. The voltage offset V0 * is added for a prompt start-up. The voltage vector position \u03b8v * is obtained by integrating the voltage vector speed \u03c9v *. The operator dvqv to \u03b1\u03b2 transforms the voltage vector vs *(V*, \u03b8v *) in polar coordinate into vs *(v\u03b1*, v\u03b2*) in \u03b1\u03b2 stator reference (Fig. 5). The voltage vector speed correction is added to prevent the synchronism losing [1], [2]. For steady state operation, the voltage vector position in dq rotor reference (\u03b8vd) is constant, therefore the voltage vector speed \u03c9v = \u03c9r because \u03c9vd = 0 (Fig. 5). In order to compensate the rotor speed variation, which appears in transient state operation, the corrected voltage vector speed reference \u03c9v * takes the form * *\u03c9 = \u03c9 + \u0394\u03c9v r r , (12) where \u0394\u03c9r is the transient rotor speed variation. The rotor speed variation \u0394\u03c9r is active only in transient state, and because it is not available for measurement (the absence of position/speed encoder), \u0394\u03c9r is estimated using the active power variation \u0394P [1], [2]. For a step load torque, the rotor speed decreases and because the voltage amplitude is initially constant (the voltage amplitude correction is not so fast), the rotor speed correction \u0394\u03c9r is obtained from active power variation \u0394P based on (1-6): /\u03c9 \u03c9\u0394\u03c9 = \u22c5\u0394 \u22c5 \u03c9\u2212 = \u0394r e rK T K P , (13) where K\u03c9=20 is a gain experimentally obtained",
            " (14) The active power variation \u0394P is extracted from P using a 1st order high pass filter (HPF), with experimentally chosen time constant T=0.2 s. The voltage amplitude correction \u0394V is based on the power factor angle regulation loop. V/f control uses the voltage vector reference frame dvqv with the dv axis oriented along the voltage vector vs. The electromagnetic torque Te is controlled by the current vector \u03d5= j s si I e using the voltage amplitude V, where \u03c6 is the angle between voltage and current vectors, i.e., the power factor angle (Fig. 5). A PI controller is used for the power factor angle regulation loop, with the voltage amplitude correction \u0394V as output. The time constant Ti= Ls/Rs \u22480.005 is chosen close to the electrical time constant, and the proportional constant kp=0.2 is chosen to avoid chattering and to obtain a desired bandwidth. The power factor angle is given by ( )\u03c6 2 ,= atan Q P , (15) where Q is the instantaneous reactive power, computed as ( )* *3 / 2 \u03b2 \u03b1 \u03b1 \u03b2= \u2212Q v i v i . (16) In order to compare the performance of the proposed sensorless V/f control with two stabilizing corrections, a Field Oriented Control (FOC) standard strategy with measured position from an encoder is used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002349_978-3-319-45781-9_26-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002349_978-3-319-45781-9_26-Figure3-1.png",
        "caption": "Fig. 3. Designed and prototyped joint samples.",
        "texts": [
            " The hole diameter is increased considering the minimum necessary clearance to have two separate movable parts after the manufacturing and post processing procedures. The test plane is conceived as represented in Table 1 and Table 2, considering pliers and mechanical arm dimensions respectively. Four specimens are designed and manufactured for each reference nominal shaft diameter; the respective holes\u2019 diameters are increased by 0.2 mm, 0.3 mm, 0.4 mm, 0.6 mm and 0.8 mm. 20 sample joints are designed and prototyped. The selected geometry and the prototyped models are represented in Figure 3. The optimal clearance values are determined as a balance of both the possibility to move the parts of each joint, with the posable factor. Since the case study has only a research purpose, the assessment is performed manually, for further de- velopment on a real case study, the real forces and strains have to be considered and applied to each specimen. The selected clearance values are than used to optimize the robotic mechanism sizing. 4 Results Fig. 4. Final prototyped model. The manufactured sample joints are used to assess the necessary clearance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000601_crv.2015.46-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000601_crv.2015.46-Figure1-1.png",
        "caption": "Figure 1. The SOI is defined by the horizontal angle \u03b8h and vertical angle \u03c8v of the FOV of the vision sensor. The radius of sphere of influence rsoi is defined by the characteristics of the vehicle.",
        "texts": [
            " A SOI is defined by the horizontal angle \u03b8h and the vertical angle \u03c8v of the Field of View (FOV) of the vision sensor(s). The radius of the SOI, given by rsoi defines the boundary between activation and deactivation of the collision avoidance system. A critical radius rm, which is the distance traveled by the vehicle under maximum deceleration, specifies the distance at which the repulsive potential goes to infinity. The defined sphere of influence provides a safety zone in which the collision avoidance system is active. The sphere of influence is shown graphically in Figure 1. The proposed repulsive potential field is analogous to the electric potential field. The magnitude is inversely proportional to the distance between the vehicle and the obstacle and is also dependent on the relative velocity between them. By Assumption 3, the relative velocity between the vehicle and any point u in the set U can be determined and is given by vvo(t) = [ vv(t)\u2212 vo(t)] \u1d40 nvo (1) where nvo is a unit vector pointing from vehicle to the obstacle, vv is the velocity of the vehicle and vo is the velocity of the obstacle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.85-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.85-1.png",
        "caption": "Figure 3.85 Resolver schematic",
        "texts": [
            " More recently, they are being used in aircraft and missile control. The earliest resolvers applied excitation to the rotor winding via a slip ring assembly and typically operated at either 60 or 400 Hz. Due to this mechanical configuration, life and speed were limited. Modern designs include a rotary transformer mounted co-axially on the rotor to supply power to the rotor winding, eliminating the deficiencies of the slip ring version (see Figure 3.84). The stator windings are physically located 90\u25e6 apart on the laminated stator assembly (see Figure 3.85). 154 Electromechanical Motion Systems: Design and Simulation The rotor has a constant AC reference voltage applied to it with the result that each stator winding has an output voltage that varies from maximum in-phase with the rotor to zero to maximum out-of-phase to zero and back to maximum in-phase. Since the stator windings are 90\u25e6 apart, their voltages are proportional to the sine and cosine of the shaft angle as it rotates through 360 mechanical degrees (one revolution). Figure 3.86 shows the waveforms for a resolver with 2 kHz excitation rotating at 3000 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002817_j.surfcoat.2020.125463-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002817_j.surfcoat.2020.125463-Figure1-1.png",
        "caption": "Fig. 1. Upskin exposure strategy of the BMG substrates.",
        "texts": [
            " While these investigations were all implemented on casted BMGs, the present state of knowledge regarding film depositions on BMGs, which are manufactured by means of LPBF-M, is strongly restricted. For this reason, it will be analyzed how the microstructure and surface topography of Zr-based LPBF-M fabricated BMG substrates correlate with the microstructural and mechanical properties of additionally applied ZrN films. Differing surface roughness conditions of the BMG substrates will be generated by means of varying upskin exposure parameters that adapt the laser power and the scan speed and correspondingly the energy-input (Fig. 1). In this context, the influence of these process strategies on the amorphous fraction within the surface layer will be examined. Furthermore, the roughness will be decreased by grinding and polishing post-treatments. The crystal structure, residual stresses as well as the hardness, Young's modulus, and the adhesion of the ZrN films in dependence of the substrate roughness and microstructural condition will be analyzed. If the films grew crystalline on metallic glasses, temperature-induced crystallizations of BMGs in tribological contacts could be avoided",
            " The machine is equipped with a Yb-fibre laser emitting a laser beam with a wavelength of 1064 nm and a maximum power of 200 W distributed over a spot size diameter of 40 \u03bcm. The layer thickness during LPBF-M processing was set to 20 \u03bcm with a rotating stripe exposure strategy for all samples. The bulk substrates were processed with a volume energy density of 25.0 J/mm3 (powder A) and 31.25 J/mm3 (powder B), respectively. According to Table 1, parameter adjustments as well as remelting processes within the last 20 \u03bcm layer of each sample were conducted to influence the surface roughness of the BMG substrates (cf. Fig. 1). Remelting was subdivided into a double exposure, subsequently applying the same energy density twice onto the last layer, and low energy remelting, applying a significantly reduced energy density by 50% within the remelting of the last layer. As one preparation step for the film deposition, the specimens S1 and S2 were mechanically post-treated by diamond grinding and polishing up to a grain size of 3 \u03bcm. All other samples (S3\u2013S9) were left in their asprocessed conditions. Prior to the film deposition, all BMG substrates were cleaned for 15 min in an ultrasonic bath filled with ethanol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003380_s1068366620040029-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003380_s1068366620040029-Figure2-1.png",
        "caption": "Fig. 2. Diagram of contact between bearing bearing with tracks of rings. Here ; , where and are deformation at contacts between bearing and rings.",
        "texts": [
            " In addition, an increase in the clearance in rolling bearings affects the stiffness of the rotor bearings and, as a consequence, the eigenfrequencies of the system 359 int ext W int ext\u0410B R R D= + \u2212 + \u03b4 + \u03b4 0 int ext W; 2iB\u041e S A\u041e R R D g= = + \u2212 \u2212 int\u03b4 ext\u03b4 [6]. Therefore, it is very important to know the true value of the radial clearance in rolling bearings, which changes during operation. Objective\u2014To develop a method for calculating the durability of bearings taking into account the wear of rolling elements and study the effect of wear on the performance characteristics of the bearings. Figure 2 shows a view of the bearing assembly. Neglecting the centrifugal forces of the bearings and possible distortions of the bearing rings, let us determine the geometric relationships in the bearing. The main notation and contact diagram of the bearing bearing with the rings are shown in Fig. 2. The initial contact angle \u03b10 in the bearing is deter- mined by the dependence (1) where g is the diametric gap; Rint and Rext are the radii of the gutters; and DW is the diameter of the bearing. Under loads, radial displacement N0 occurs in the bearing, and axial displacement S0 of the ring. Due to the small radial displacement, the size of the diametric gap in the section passing through the center of the bearing is determined by the formula (2) where is the angular coordinate of the bearing, measured from the line of action of the radial load; z is the number of bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure21-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure21-1.png",
        "caption": "Fig. 21 Total deformation (in lm) at final load step",
        "texts": [
            " That is, the socket and the skirt do not need to deform much beyond the point at which the campaniform sensilla are activated, so the hair begins to slide to distribute the stresses more evenly through the structure. 5.5 Socket Spring Constant. Calculation of the torsional spring constant of the socket is important to understand the dynamic behavior of the hair\u2013socket assembly. The spring constant can be calculated using Sh(t)\u00bc T(t), where h is the angular displacement of the hair, T is the applied torque on the hair shaft, and S is the spring constant of the assembly. To find the torque, the center of rotation had to be determined. This was calculated very easily by studying the total deformation plot (Fig. 21) and matching the contour bands on the plot to the legend. The center of rotation of the hair occurs at 29.3 lm from the bottom of the hair base. This point also exactly lines up with the top edge of the belt. This makes sense because the belt compresses on one side and tenses on the other, and its tension/compression pulls the hair with it (of course, the socket base still mediates the hair movement). The spring constant, like other parameters calculated for the model, can be considered as having two distinct value regimes, corresponding to the movement ranges before and after iris contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002571_2168-9806.1000135-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002571_2168-9806.1000135-Figure1-1.png",
        "caption": "Figure 1: 4100C Boss Electric Shovel [1].",
        "texts": [
            " This is an open-access article distributed under the terms of the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original author and source are credited. Keywords: Surface mining; Crawler-terrain interactions; Multibody dynamic theory; Crawler contact forces; Terrain deformation; Virtual prototype simulation Cable shovels are widely used in surface mining operations. The lower works of this shovel comprise propel and crawler systems, which support the upper body and attachment as shown in Figure 1 [1]. The crawler tracks are made up of crawler shoes that are connected together by link pins to form a continuous chain [2]. The multi-body dynamics study on the interaction between the shovel crawler track and the terrain for large shovels in surface mining operations is not widely available in the open literature. This study is important to determine crawler shoe kinematics, contact forces between crawler tracks and ground and reaction forces at the link pin joint between adjacent crawler shoes",
            " [11] developed virtual prototype simulators in MSC ADAMS to examine the GAP (ground articulating pipeline) torque requirements and GAP carriage-oil sand terrain interactions. They used a contact force model similar to that in [10] to account for GAP track-oil sands interactions. They also used flexible oil sand model to illustrate the dynamic deformation of oil sand terrain under GAP carriage motion. Frimpong and Thiruvengadam [12,13] formulated the kinematic and dynamic equations of motion that govern the two types of propelling motion of the shovel crawler track (P&H 4100C BOSS Electric Shovel in Figure 1) on flexible oil sand terrain based on multibody dynamics theory. Their study developed the 3-D virtual prototype simulator of crawler track interacting with oil sands terrain within MSC ADAMS and reported the simulation results for crawler track shoe kinematic quantities (linear and angular displacement, velocity and accelerations). In this paper the time varying contact forces that are used as an input to calculate crawler shoe kinematic quantities reported in Frimpong and Thiruvengadam [12,13] are presented",
            " One link pin is made a spherical joint and the other link pin is made a parallel primitive joint to create equivalent revolute joint between two crawler shoes. The crawler shoe model is generated in Solidworks based on the actual crawler shoe model for P&H 4100C Boss shovel [14]. Similarly, the oil sand terrain is made up of spring- damper-oil sand units connected to four adjacent oil sands units by spherical joints [10]. The stiffness (k) and damping (c) values of oil sand terrain are listed in Table 2. Only the open track chain of the crawler assembly, in contact with the ground Figure 1, is used for this study. More details on the dimensions, joints and material properties of crawler shoe and oil sand units can be found in Frimpong and Thiruvengadam [12,13]. The global coordinate system is located at the left corner of the oil sand terrain at point O as shown in Figure 2. The position of the center of mass of crawler shoes 1 - 13 and oil sand units 15 and 64 at time t = 0 with respect to the global coordinate system are listed in Table 3. The joint locations of spherical and parallel primitive joints between each crawler shoes are listed in Table 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure12-1.png",
        "caption": "Figure 12. Simplified schematic diagram of the size comparison between deformation and profile interference.",
        "texts": [
            " If a suitable y is found, the maximum contact stress Pmax under the actual normal load Qw can be obtained by Pmax \u00bc max\u00bdP\u00f01\u00de,P\u00f02\u00de, . . . ,P\u00f0n\u00de \u00f025\u00de Correction for error caused by profile geometry interference. When a point contact occurs between the wheel and rail surfaces, the contact can be described by an equivalent model, in which a curved surface is in contact with a flat surface. If replacing the deformation with profile geometry interference, errors will be introduced in the sizes of the interference region and actual contact deformation zone when the interference value is equal to the amount of deformation. Figure 12 shows a simplified schematic diagram of the error in the length direction of the contact spot, where apc is the size of the actual contact deformation zone; asl is the size of the interference region; pc is the deformation amount; and sl is the interference value. In calculating the wheel/rail stress, it is necessary to correct these errors. For the deformation of a point contact, the following relationships can be written apc Rz \u00bc CA W1=3 PC \u00f026\u00de W2=3 PC \u00bc pc CD Rz \u00f027\u00de Combining equations (26) and (27), we can obtain equation (28) apc Rz \u00bc CA ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pc CD Rz r \u00f028\u00de However, when wheel/rail profile interference occurs, sl can be obtained by sl a2sl 2 Rx \u00f029\u00de Assuming pc\u00bc sl, equation (30) can be obtained by combining equations (28) and (29) asl apc \u00bc 1 CA ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 k CD p \u00f030\u00de For different k, there must be a proportional relationship between asl and apc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003476_0954408920948683-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003476_0954408920948683-Figure4-1.png",
        "caption": "Figure 4. Kinematics nomenclature of picking robot.",
        "texts": [
            " Using the arc length theorem and the trigonometric function, the following results can be obtained DL1 \u00bc h h1 Sin\u00f0a /\u00de DL2 \u00bc h h2 Sin\u00f0a\u00fe b /\u00de DL3 \u00bc h h3 Sin 3p 2 / 8>>< >>: (2) Thus, the drive of the continuum structure can be achieved by controlling the variation of the wire by the servo motor. In order to obtain the workspace of the picking robot, the transformation matrix is used to analyze the forward kinematics solution of the picking robot. The calibration of the coordinate system of the picking robot is shown in Figure 4. Coordinate system x\u03021; y\u03021; z\u03021 is the original coordinate system, point 1 is the midpoint of the turntable. Coordinate system x\u03022; y\u03022; z\u03022 is the rigid frame part end coordinate system, point 2 is the end point of rigid frame. Coordinate system x\u03023; y\u03023; z\u03023 is the continuum part base coordinate system, point 3 is the center of the first tendon guide of continuum structure. Coordinate system x\u03024; y\u03024; z\u03024 is the terminal coordinate system, point 4 is the center of the top tendon guide of continuum structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001431_s10846-014-0044-7-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001431_s10846-014-0044-7-Figure1-1.png",
        "caption": "Fig. 1 Structure of MMAR [3]",
        "texts": [
            " Such flying-walking locomotion greatly improve the activities of birds. For robot development, the flying-walking locomotion is obviously attractive. The recent advanced researches focus on aerial robots coincide with the demand for flying-walking locomotion. Obviously, the flying-walking locomotion must be executed by some novel flying mechanisms. Inspired by the merits of the quadrotor, in our previous research, a novel multi-propeller multifunction aerial robot (MMAR) has been presented as shown in Fig. 1 [3]. This robot is composed of a quadrotor main-body subsystem and manipulator subsystem. Each manipulator has two joints. The four joints can rotate in a plane that is normal to the propellers\u2019 plane. The two manipulators can be used as legs or arms so as to advance the function of the aerial robot. Such a novel design of MMAR make it possible to execute the flying-walking locomotion. To achieve the flying-walking locomotion, the MMAR must land on the ground before it transitions from flight to walking"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.25-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.25-1.png",
        "caption": "Fig. 18.25 Schematic diagram of chemical thermal sensor",
        "texts": [
            " Therefore, a chemical reaction, which is associated with heat, can be detected by an appropriate thermal sensor, such as described in Chap. 17. These sensors operate on the basic principles that form the foundation of a microcalorimetry. The operating principle of a thermal sensor is rather simple: a temperature probe is coated with a chemically selective layer. Upon a chemical exposure, the probe measures transfer of heat during the reaction between the sample and the coating. A simplified drawing of such a sensor is shown in Fig. 18.25. It contains a thermal shield to reduce heat loss to the environment and a thermistor coated by a catalytic layer. A biosensor based microcalorimeter may be made where the sensitive layer has an enzyme immobilized into a matrix. An example of such a sensor is the enzyme thermistor covered by an immobilized glucose oxidase (GOD). The enzymes are immobilized on the tip of the thermistor, which is then enclosed in a glass jacket in order to reduce heat loss to the surrounding solution. Another similar sensor with similarly immobilized bovine serum albumin is used as a reference"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003305_s10035-020-01039-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003305_s10035-020-01039-5-Figure1-1.png",
        "caption": "Fig. 1 Experimental method. a Photo of the experimental set-up taken at the end of a test. b Illustration of the initial position of the plate in the glass bead container (distances are expressed in milimeters)",
        "texts": [
            "\u00a03, we empirically characterise the drag fluctuations by quantifying their typical frequency and magnitude. In Sect.\u00a04, we * P. Rognon pierre.rognon@sydney.edu.au 1 School of\u00a0Civil Engineering, The University of\u00a0Sydney, Sydney, NSW\u00a02006, Australia T.\u00a0Hossain, P.\u00a0Rognon 1 3 72 Page 2 of 17 introduce a physical process that can explain the observed properties of this drag instability. This section presents the experimental method used to conduct vertical uplift tests and measure the resulting drag forces. Uplift tests were conducted using the experimental set-up presented in Fig.\u00a01. This set-up has previously been used in [23] to measure the uplift capacity of anchors with a shape mimicking tree roots. It is comprised of a container filled with glass beads, in which a PDMS disk-shape plate is buried at a designated depth. The plate\u2019s motion is driven by a loading frame (H5KS Olsen Loading Frame) via a stainless steel shaft of diameter 4\u00a0mm. The plate has a circular cross section of diameter B = 40\u00a0mm, and is 4\u00a0mm thick. The container diameter is 170\u00a0mm, which is about four times larger than the plate",
            " We attribute this relatively low porosity to the possible existence of fine grains in the packing and a possible distribution in sphericity of the beads. The preparation of the uplift tests involves filling the first 80\u00a0mm of the container with glass beads, placing the plate horizontally at this location, and adding glass beads above it up to the desired level. All tests presented in the following are performed with an initial plate depth of H=120\u00a0mm corresponding to an embedment ratio H B = 3 . Unless otherwise specified, the plate is placed at an equal distance from the vertical edge of the container (see Fig.\u00a01b). After this preparation, the uplift test involves vertically pulling the plate at a prescribed constant velocity v. The loading frame drives this motion using a DC servo motor and records over time the plate\u2019s vertical displacement P(t) = vt as well as the drag force F required to achieve it. The frequency of force and displacement readings is set to 100\u00a0Hz. The force is zeroed during the preparation just before the plate is covered by grains. This means F measures the net reaction of the granular packing on the plate and excludes the self-weight of the plate and shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure4.18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure4.18-1.png",
        "caption": "FIGURE 4.18 Capacitor with two different connecting leads.",
        "texts": [
            " At xoffset = 3, we can't see the difference between the h = 1.0 and the h = 0.1 cases because, again, each electrode no longer knows where the other one is. Figure 4.16 shows another possible variation. In this case both electrodes are kept square and centered, but the size of one of the electrodes is varied. Figure 4.17 shows the results of these variations when h = 0.1. There is no reason why either of the electrodes has to be square, or why different-size electrodes and misorientation cannot be combined. Figure 4.18 shows a capacitor with connecting leads brought out. In this example the electrodes are rectangular and the two connecting leads do not have the same dimensions. 4.8 Varying the Geometry 83 Table 4.3 shows the data file for mom1.m that creates this capacitor. The format of the file is the same as in the previous examples. All dimensions are inmeters. The first line in the data file is the cell 12 side, a = b = 0.01. The second line describes the electrode of the lower section. The third line describes the connecting lead to this (lower) electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000992_s11044-013-9388-1-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000992_s11044-013-9388-1-Figure8-1.png",
        "caption": "Fig. 8 Inverted double pendulum model",
        "texts": [
            " These combinations and partitions are repeated while there is a pair of parent node and child node that will decrease \u0394S\u2217. Note that this operation in the while loop does not change the total number of cells. Systems need some closed loop control method that resists against disturbances to realize the planned motion of RASMO. A state feed\u2013back control method is employed that randomly samples predicted future states as in [22]. Motion planning simulations are applied to a inverted double pendulum model (Fig. 8) to validate the applicability of the proposed partitioning method. The proposed method is compared with a uniform partitioning method in this experiment. 3.1 Settings The motion equation for the pendulum in Fig. 8 is Aq\u0308 + B = \u03c4 (12) where A = [ \u03b1 + \u03b3 + 2\u03b2 cos(q2) \u03b3 + \u03b2 cos(q2) \u03b3 + \u03b2 cos(q2) \u03b3 ] (13) B = [\u2212\u03b2(2q\u03071 + q\u03072)q\u03072 sin(q2) + k1 cos(q1) + k2 cos(q1 + q2) \u03b2q\u03072 1 sin(q2) + k2 cos(q1 + q2) ] (14) \u03b1 = J1 + 1 4 m1l 2 1 + m2l 2 1 , \u03b2 = 1 2 m2l1l2, \u03b3 = J2 + 1 4 m2l 2 2 , k1 = ( 1 2 m1l1 + m2l1 ) g, k2 = 1 2 m2l2g. (15) Table 2 Parameters of pendulum model Joint Length l [m] Mass m [kg] Center of mass [m] 1 0.5 0.5 0.25 2 0.5 0.5 0.25 Table 3 Parameters of planning system N.D. means nondimensional Joint Position limitation [rad] Velocity limitation [rad/s] Maximum torque [N m] Torque search resolution [N"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.22-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.22-1.png",
        "caption": "Fig. 18.22 Concept of the measurement setup for optical cantilevers (adapted from ref. [53])",
        "texts": [
            " The bending of the microcantilever is not caused by the weight of the absorbed chemical but by the absorption induced surface stresses due to changes in surface free energy. The cantilever will bend if the surface free-energy density change is comparable with the cantilevers spring constant. When chemicals come in contact with a coated cantilever, electrostatic repulsions, swelling, or other affects result in changes in surface stress, which ultimately result in measurable cantilever bending. Cantilevers can be measured in many ways. Originally the sensing systems were based on optical (laser) detection (Fig. 18.22) being developed for SPM. Newer research has implemented thermal [65], capacitive [41], and piezoresistive [68] measurement techniques, thereby removing the need for lasers and associated optics, resulting in a simpler measurement circuit. Spectrometry (spectroscopy) is a class of powerful methods used to analyze chemical compositions based on energy or mass. There are several types of spectrometry, including: Mass Spectrometry (MS): Sample molecules are ionized, and the mass to charge ratio of these ions is measured very accurately by electrostatic acceleration and magnetic field perturbation, providing a precise molecular weight"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure11-1.png",
        "caption": "Fig. 11 The bottom faces of the hair base and socket base were constrained to not move along the short (x) axis only",
        "texts": [
            "org/about-asme/terms-of-use purposes. The skirt is modeled with uniform thickness to avoid further complexity. It is apparent that nonuniformity in thickness will affect local bending of a plate. As a first analysis, we are interested in the global deformation of the skirt and it is assumed that a small variation of skirt thickness will not have significant effect on the global deformation behavior of the skirt. The bottom faces of the hair base and socket base were constrained by a displacement condition (Fig. 11). If the displacement is defined by the Cartesian components of a face, then a nonzero input for a given direction means that the face will retain its shape but move relative to its original location, based on the magnitude and direction of the force. The x-direction displacement (in and out of the viewing plane) was set to zero, and the other two directions were left free to move, rotate, or deform. This represents the actual motion constraint in the insect, as the hair moves in zdirection (perpendicular to the long axis), and the displacement in x direction (along the long axis direction) is considered negligible (and set to zero in the FEA setup)",
            " A similar process occurred on the other side (opposite to the hair shaft deflection), but in reverse: the lower half of the hair base slid out of contact, though the separation was small compared to the case described. This is probably because the lower part of the socket base beyond the diagonal edge is thicker and provides more support. The second result was that there was some sliding of the hair base on the inside socket cavity wall along the long or major axis of the elliptical cross section of the geometry (x-axis in Fig. 11). Based on the observations from the contact status above, a higher pressure value was expected between the contacts in the hair movement direction near the bottom of the socket cavity than for values opposite the hair movement direction near the top of the socket cavity, because the gap was smaller at the socket base. This is exactly what was seen, with the highest contact pressure of 40 kPa (Fig. 14) and the lower value on the other side around 12 kPa. Along the long axis where the sliding occurred, pressures of up to 22 kPa were observed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002237_978-3-319-29357-8_59-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002237_978-3-319-29357-8_59-Figure4-1.png",
        "caption": "Fig. 4 The example cases when the grasp type changes during the manipulation due to low friction coefficient: a, b pinch grasp changes into spherical\u2013like grasp and the orientation of the object changes, c, d pinch grasp changes into power\u2013like grasp as the object slips towards the palm during the fingers closure",
        "texts": [
            " Execution of each selected grasp was performed in dynamics simulation in two steps: \u2219 perform the grasp: close the fingers on the object, \u2219 lift-up: move the gripper against the gravity force. The grasp was considered successful if the object could be lifted and held by the gripper above the specified height above the ground plane. The results of the grasping simulation are presented in Table 2. The Coulomb friction coefficient was set to relatively low value \ud835\udf07 = 0.1 during all experiments. This caused that only the most stable grasps were successful. The low friction force caused the change of some grasps types during the finger closure (Fig. 4) and (Table 3). Ta bl e 3 E x a m p le s o f s u c c e s s f u l g ra s p s g e n e r a te d fo r v a r io u s o b je c ts a n d g ra s p m o d e ls O b je c t G r a s p m o d e l P in c h S p h e r ic a l P o w e r H o o k C o n ta in e r 1 C o n ta in e r 2 U n li k e ly C u b o id U n li k e ly U n li k e ly C y li n d e r ( ly in g ) U n li k e ly U n li k e ly ( c o n ti n u e d ) Ta bl e 3 ( c o n ti n u e d ) O b je c t G r a s p m o d e l P in c h S p h e r ic a l P o w e r H o o k C y li n d e r ( s ta n d in g ) J a r U n li k e ly U n li k e ly K e tt le M u g U n li k e ly U n li k e ly The method presented in this paper is a modification of [1]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure17-1.png",
        "caption": "Fig. 17. Non-orthogonal gear pair.",
        "texts": [
            " This third example considers a special case of non-orthogonal axes gearing with a small shaft offset to further illustrate the use of evolutoids and involutoids to define conjugate gear flanks. The selected example is based in part on an existing non-orthogonal spiral bevel gear set [23] . The shaft angle and offset are changed to establish an unconventional hypoid gear set, one where the input pitch surface is conical and the output pitch surface is planar and perpendicular to its rotation axis. A summary of the non-orthogonal gear data is presented in Table 4 . Depicted in Fig. 17 is a virtual model of the non-orthogonal gear pair. The axial pitch of the ISA (Instantaneous Screw Axis) is zero when the shaft offset E is zero, and as a result, the variation in pressure angle is unconstrained for pure bevel gears. In this example, the variation in normal pressure angle is based on general hypoid gears in mesh (where E = 5 . 00 mm) [2] . The required variations in spiral angle and pressure angle are presented in Fig. 18 . The nominal contact pattern for the unmodified flank data (no transverse or axial profile modification) is provided in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001392_s11721-013-0084-9-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001392_s11721-013-0084-9-Figure1-1.png",
        "caption": "Fig. 1 (a) Zones of repulsion, alignment and attraction. The radius and thickness of each zone is determined by dj and mj , respectively, where j = r,al, a. (b) Generic contour plot of the attractive distance kernel Kd a",
        "texts": [
            " Factors of influence The model depicts an individual\u2019s decision to continue moving in its current direction or to turn based on its neighbors. We assume that the only two factors that influence turning are the distance from neighbors and the neighbors\u2019 orientation. Each factor is modeled mathematically by interaction kernels. (i) Distance kernels. Each individual is assumed to be surrounded by zones of repulsion, alignment, and attraction acting at short, intermediate and long ranges, respectively (see Fig. 1(a)). These zones are modeled with distance kernels Kd j : Kd j ( x) = 1 Aj e \u2212( \u221a x2+y2\u2212dj )2/m2 j , j = r,al, a, (3) where j = r,al, a, stands for repulsion, alignment, or attraction, respectively, dj represents the radius and mj the thickness of the respective influence zone (see Fig. 1(a)). The Aj \u2019s are normalizing constants that make the spatial integral of each kernel equal to 1. These constants are given by Aj = \u03c0mj ( mje \u2212d2 j /m2 j + \u221a \u03c0dj ( 1 + erf(dj /mj ) )) . (4) Figure 1(b) shows the attraction kernel Kd a . Note that the three interaction zones may overlap. (ii) Orientation kernels. We discuss first the attractive orientation kernel. Suppose a decision making individual located at x is heading in direction \u03c6 and senses a neighbor located at s, within its attraction zone\u2014see Fig. 2(a). The relative location s \u2212 x makes an angle \u03c8 with the positive x axis. Due to attraction, the reference individual makes a decision to turn in order to approach its neighbor at s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001435_iecon.2013.6699639-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001435_iecon.2013.6699639-Figure7-1.png",
        "caption": "Fig. 7. For a healthy rotor: a) Magnetic flux density; b) Current density distribution.",
        "texts": [
            " As the number of broken bars increases, the fluctuations worsen and the average torque decreases. Forces affecting each bar were calculated using the same FEM software that was used to calculate the inductances. For all cases the magnetic flux density distribution, the current density distribution and the tangential and radial forces are calculated at an instant where the current in phase \u2018a\u2019 is maximum. The magnetic flux density and current density distributions for a healthy rotor are shown in Fig. 7. The radial and tangential forces are shown in Fig. 8. It can be seen from these figures that, for the case of healthy rotor, the radial forces are almost symmetric and cancel each other. On the other hand, the tangential forces, which are responsible for generating the motor torque, are symmetrical under the machine poles. The magnetic flux density and current density distributions for the case of one broken bar are shown in Fig. 9. Distributions of the tangential and radial forces corresponding to the same instant are shown in Fig",
            " It can be observed from Fig. 12(a) that the asymmetry in the radial forces has increased in comparison to the case of one broken bar. Such increase in the asymmetry would negatively affect the bearings life time. Increased asymmetry of the tangential forces, as illustrated by Fig. 12(b), would result in increased torque and speed variations. The negative tangential forces cause a decrease in the average torque compared to the healthy case. It can be seen from the magnetic flux density distributions, shown in Fig. 7(a), Fig. 9(a) and Fig. 11(a) that the distribution in the faulty cases are asymmetric compared to the healthy case. This paper presents an efficient model to analyze broken bar fault conditions for induction motors using a FEM approach coupled with an ABC transient model. The pattern of the asymmetry in bar currents resulted from various broken bar faults can be deduced and analyzed. Effects on the motor torque and speed can be observed. Radial and tangential force asymmetries resulting from broken bars conditions can be assessed using the proposed approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001508_0954406214553983-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001508_0954406214553983-Figure7-1.png",
        "caption": "Figure 7. Application.",
        "texts": [
            " To solve a problem of collision without friction of two rigid bodies, one may use the following algorithm: \u2013 calculate the parameters dj, ej, fj dj \u00bc YjAcj ZjAbj, ej \u00bc ZjAaj XjAcj, fj \u00bc XjAbj YjAaj \u00f052\u00de \u2013 calculate the velocities v01n, v 0 2n v01n \u00bc N1f g T g\u00bd v01 , v02n \u00bc N2f g T g\u00bd v02 \u00f053\u00de \u2013 calculate the inertances using equation (44) and the intensity of the impulse using the relation P \u00bc 1\u00fe k\u00f0 \u00dev012n g1 \u00fe g2 \u00f054\u00de \u2013 calculate the velocities after the collision using equation (46); \u2013 calculate the loss of kinetic energy Ec \u00bc 1 k2 v012n 2 2 g1 \u00fe g2\u00f0 \u00de \u00f055\u00de Application Two rectangular homogeneous shells of dimensions 2li, 2hi, i \u00bc 1, 2, and masses mi, i \u00bc 1, 2 collide (Figure 7) at point A. Knowing the velocities w1, w2 of the centers of weight C1, C2, the angle , and the distance BA \u00bc s, we determine the impulse P, the velocities after the collision, and the loss of kinetic energy. Numerical application is as follows: m1 \u00bc 800 kg, m2 \u00bc 1000 kg, l1 \u00bc 2m, h1 \u00bc 0:75m, l2 \u00bc 2:5m, h2 \u00bc 1m, s \u00bc 2m, \u00bc 30 , k \u00bc 0:4, w1 \u00bc 20m=s, w2 \u00bc 30m=s. Solution: From Figure 7, the followings are resulted a1 \u00bc sin , b1 \u00bc cos , c1 \u00bc 0, d1 \u00bc 0, e1 \u00bc 0, f1 \u00bc l1 cos h1 sin , a2 \u00bc 0, b2 \u00bc 1, c2 \u00bc 0, d2 \u00bc 0, e2 \u00bc 0, f2 \u00bc s, X1A \u00bc l1, Y1A \u00bc h1, X2A \u00bc s, Y2A \u00bc h2 and one obtains the column matrices N1f g \u00bc sin cos 0 0 0 f1 T \u00bc sin cos 0 0 0 l1 cos h1 sin T N2f g \u00bc 0 1 0 0 0 s T v01 \u00bc 0 0 0 w1 0 0 T , v02 \u00bc 0 0 0 w2 0 0 T It also results in J1z \u00bc m1 l21 \u00fe h21 12 , J2z \u00bc m2 l22 \u00fe h22 12 g1 \u00bc 1 m1 \u00fe f21 J1z , g2 \u00bc 1 m2 \u00fe f22 J2z v01n \u00bc N1f g T g\u00bd v01 \u00bc w1 sin , v02n \u00bc N2f g T g\u00bd v02 \u00bc 0 v012n \u00bc w1 sin P \u00bc 1\u00fe k\u00f0 \u00dev012n g1 \u00fe g2 M1C\u00bd 1 N1f g \u00bc 0 0 f1 J1z sin m1 cos m1 0 T M2C\u00bd 1 N2f g \u00bc 0 0 s J2z 0 1 m2 0 T v1f g \u00bc v01 P M1C\u00bd 1 N1f g, v2f g \u00bc v02 \u00fe P M2C\u00bd 1 N2f g Ec \u00bc 1 k2 v012n 2 2 g1 \u00fe g2\u00f0 \u00de at University of Sydney on October 7, 2015pic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.1-1.png",
        "caption": "Figure 4.1 Graphical example of accuracy and repeatability",
        "texts": [
            " Repeatability is a measure of how well a system returns to a specific position over a number of identical moves. It is a statistical quality gathered over those moves. It can be specified as unidirectional or bidirectional. Unidirectional is easier to achieve since backlash does not affect the operation, but motion to the commanded position must always occur from the same direction, increasing total motion time. Bidirectional requires close control of backlash, pre-loads, belt/pulley/coupling compliance, and so on. Figure 4.1 shows a graphical example of accuracy and repeatability. Velocity expresses the speed of the load, or the rate at which the load is changing its position. 170 Electromechanical Motion Systems: Design and Simulation Specifying the maximum system velocity and in turn the maximum motor velocity, together with the anticipated supply voltage will contribute directly to determining the voltage constant of the motor (Ke). This then results in the motor torque constant (Kt), which together with the acceleration requirement will determine the peak motor current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure20-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure20-1.png",
        "caption": "Figure 20. Speed field contours for 500 rpm.",
        "texts": [],
        "surrounding_texts": [
            "Table 2 and the graph (Figure 19) show the calculations of the maximum speed of the velocity fields for a fixed drum diameter of 650 mm and different values of the drum rotation speeds. The diameter and rotation speed of the beater are equal to the initial values. [22] ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0003733_icems50442.2020.9290866-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003733_icems50442.2020.9290866-Figure1-1.png",
        "caption": "Fig. 1. Proposed motor",
        "texts": [
            " In this paper, detail structure and basic theory of the proposed motor are denoted in section II. In section III, numerical simulations and experimental results using prototype motor are shown. The results contain the reduction effect of weakening current in the proposed method. Moreover the validation, in which between numerical simulation and experimental results are shown, is discussed. Defined parameters used in this paper is summarized in Table I. A structure of the proposed motor is shown in Fig. 1. From Fig. 1(a), it can be seen that brush and slip ring is equipped on shaft for electric power supplying to field coil. Thanks to this, magnetomotive force of the field coil can be changed actively. The electrical connection between field coil and inverter is secured by welded jointing on terminal. In Fig. 1(b), rotor assy of the proposed motor is shown. The rotor consists of laminated core and field winding unit. The field winding units, which consist of small size core, field coil and permanent magnet, are inserted from axis direction of rotor core. Since coil winding is separated from laminated rotor core, coil winding process becomes easier. It is assumed that field winding unit is fixed by using adhesive bond to laminated core same as permanent magnet. The current direction of the proposed motor is limited for assisting the magnetic flux"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003592_icmica48462.2020.9242886-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003592_icmica48462.2020.9242886-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of Axis and Coordinates system",
        "texts": [
            " Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 05:54:00 UTC from IEEE Xplore. Restrictions apply. Thus Newton-Euler formulation is used to describe the forces that act on the quadcopter [11]-[12]. In order to represent a mathematical model for the quadcopter some equations must be derived to eventually formulate the equations of motion which will help in simulating the quadcopter and add a controller to the system. Axis and coordinates system of the quadcopter with its body frame and earth frame are shown in Fig. 1 For the quadcopter are chosen the input control as U = [U1 U2 U3 U4]T where the first control input is for the altitude control while the remaining three control input are for controlling the attitude Roll (\u2205 ,Yaw ( and Pitch (\u03a8). 1. Thrust force [U1]. 2. Rolling moment (Torques) [U2, U3 and U4]. As shown in equations (1), (2), (3) and (4) [1], [13].The basic need is to stabilize the quadcopter at the required position and follow a desired reference trajectory. This is done by calculating the required rotors angular speed using inverse dynamics for the quadcopter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000833_icicip.2013.6568189-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000833_icicip.2013.6568189-Figure1-1.png",
        "caption": "Fig. 1. Loss function",
        "texts": [
            " Then, the problem of nonlinear regression in the lower dimension space is transformed into a linear regression problem in a high dimensional feature space. Hence, the optimization problem aims at searching the flattest linear function for it is a linear relationship between the input and output in the high dimensional linear feature space. For SVR, the training error is measured by means of a \u03b5 -insensitive loss function defined as follows\uff1a 0, - , if otherwise\u03b5 \u03be \u03b5 \u03be \u03be \u03b5 \u23a7 <\u23aa= \u23a8 \u23aa\u23a9 (2) Where \u03be is the slack variable which describes the deviation from the boundaries of the \u03b5 -insensitive zone illustrated in Fig. 1 where the loss function is also depicted. In order to prevent over-fitting and improve the generalization capability, the following optimization problem needs to be solved: 2 * 1 * * min 2 ( ) - , ( ) - , ( ) . 0 0 N i i i i i i i i i C y x y x s t \u03c9 \u03be \u03be \u03c9 \u03b5 \u03be \u03c9 \u03b5 \u03be \u03be \u03be = + + \u2329 \u03a6 \u232a \u2264 +\u23a7 \u23aa + \u2329 \u03a6 \u232a \u2264 +\u23aa \u23a8 \u2265\u23aa \u23aa \u2265\u23a9 \u2211 (3) where C is the regularization constant used to penalize the errors larger than \u03b5\u00b1 using -\u03b5 insensitive loss function. By using Lagrangian multipliers and Karush-Kuhn-Tucker conditions to the Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003042_j.matpr.2020.04.082-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003042_j.matpr.2020.04.082-Figure4-1.png",
        "caption": "Fig. 4. Flux lines plot of configuration 2 for I = 15A.",
        "texts": [
            "Six SOMALOY 1000 3P, 800 Mpa, pre-form blanks were procured three blanks when glued together formed the material for stator fabrication while the remaining three blanks when glued together formed the material for rotor fabrication. The properties of SMC material are sourced from Hoganas AB of Sweden while procuring the prefrom SMC blanks from which the B-H curves were drawn.The BH curves of M19 steel and SMCmaterials are shown in Fig. 3, which reveals that although the SMC has inferior relative permeability when compared with lamination steel it still posses the following desirable characteristics [4] (see Fig. 4, Fig. 5). 1. High power density by 3D magnetic flux conduction 2. Lower core losses at elevated frequencies in comparison with electrical steel 3. Good formability; complex shapes can be directly compacted without destroying the material structure and resulting deterioration of magnetic properties. The static and dynamic electromagnetic characteristics [1,5] have been obtained through FEA. MagNet 6.1.1 Finite Element Analysis software package has been employed for this while for thermal and vibration analysis ANSYS 10, another multi-physics FEA software package has been used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003534_uralcon49858.2020.9216249-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003534_uralcon49858.2020.9216249-Figure1-1.png",
        "caption": "Fig. 1. Backlash type characteristic",
        "texts": [
            " A separate equation describes the diesel engine governor, which is represented by an aperiodic link of the first order [16, 21] dhT K U h dt\u03c9 \u03c9 \u03b5= \u2212 , where is the T\u03c9 - time constant of the actuator, K\u03c9 - is the gain of the governor, U\u03b5 - is the mismatch signal between the preset 0r\u03c9 and actual frequencies r\u03c9 of diesel engine rotation. The main objective of the research is the following question: how does the phenomenon of backlash in the control circuit of the engine speed effects on the parameters of the autonomous electrical system [16]. Therefore, we take its gap Dn as the main characteristic of the backlash (Fig. 1) and use the following mathematical description to solve the problem. 401 Authorized licensed use limited to: Carleton University. Downloaded on November 03,2020 at 16:10:06 UTC from IEEE Xplore. Restrictions apply. \u0434\u043b\u044f U -k dU \u0434\u043b\u044f 0 dt n n U const D U dUk D sign dt \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 \u03b5  = \u2264    =    \u2212 \u2260         where k - is the transfer ratio, nD - is the backlash gap, \u03b5 - is the mismatch between the preset 0r\u03c9 and actual r\u03c9 diesel speeds. The mathematical model allows us to carry out the necessary research for an autonomous electrical system, in which two parallel-connected diesel-generator units operate into a common active-inductive load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002632_tencon.2016.7848445-Figure19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002632_tencon.2016.7848445-Figure19-1.png",
        "caption": "Fig. 19: Different designs of 3D printed air core inductors[22]",
        "texts": [
            " As power density and temperature requirement for the power converters are major areas to focus for the current applications, this method of building power converters will help to improve the power density. 1) Additive manufactured air core inductors: This reference [22] is a very good example of the importance of mechanical design to improve the performance of electrical passive components. AM and molding techniques have been utilized to create various types of air core inductors. AM allows the possibility to design and appreciate new design features in our prototype models. So this research had focussed on designing various shapes of air core inductors as shown in Fig. 19, and printed using polymers. Then the polymer 2016 IEEE Region 10 Conference (TENCON) \u2014 Proceedings of the International Conference 2331 printed air core inductors are used as mould to make out the the desired shapes out of metal. Moreover it\u2019s proved that some of these designs can lead to improved electrical performance. The reference also describes the design tools used by the authors to design, fabricate and characterize the electromagnetic performance of the air core inductors. They implemented 70W prototype 27"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001088_s11044-014-9417-8-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001088_s11044-014-9417-8-Figure16-1.png",
        "caption": "Fig. 16 Care must be taken so as not to arbitrarily specify constraint forces in such a manner as to violate force equilibrium. In this example, normal constraint forces applied by the robot hands against the object need to balance under static equilibrium",
        "texts": [
            " (100) We then have n = 18, mC = 12, p = 6, m = 6, N = 0, k = 12, where we will control the task as well as the normal hand constraint forces on the object and 4 of the hand constraint moments on the object, so A = \u239b \u239c\u239c \u239c\u239c \u239c\u239c \u239d 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 \u239e \u239f\u239f \u239f\u239f \u239f\u239f \u23a0 , d = \u239b \u239c\u239c \u239c\u239c \u239c\u239c \u239d \u03bb1 \u03bb4 0 0 0 0 \u239e \u239f\u239f \u239f\u239f \u239f\u239f \u23a0 . (101) Note we have chosen to specify zero constraint moments. Our system can be expressed as \u239b \u239d 1 T 0 A Sp 0 \u239e \u23a0 ( \u03c4 \u03bb ) = \u239b \u239d \u0302 T JT (\u0302cx\u0308 + \u03bc\u0302c + p\u0302c) + T (\u0302\u03b1 + \u03c1\u0302) d 0 \u239e \u23a0 , (102) where \u239b \u239d 1 T 0 A Sp 0 \u239e \u23a0 \u2208 R 30\u00d730. (103) For the grasping posture shown in Fig. 16, Rank \u23a1 \u23a3 \u239b \u239d 1 T 0 A Sp 0 \u239e \u23a0 \u23a4 \u23a6 = 29. (104) So, the system is rank deficient. However, for particular choices of x\u0308, \u03bb1, and \u03bb4 control solutions can be found. For example, given x\u0308 = 0 (static equilibrium of the object) the condition \u03bb1 = \u2212\u03bb4 produces a solution. That is, the vector \u239b \u239d \u0302 T JT (\u0302cx\u0308 + \u03bc\u0302c + p\u0302c) + T (\u0302\u03b1 + \u03c1\u0302) d 0 \u239e \u23a0 (105) can be expressed as a linear combination of the columns of \u239b \u239d 1 T 0 A Sp 0 \u239e \u23a0 , (106) when x\u0308 = 0 and \u03bb1 = \u2212\u03bb4. However, (105) cannot be expressed as a linear combination of the columns of (106) when x\u0308 = 0 and \u03bb1 = \u2212\u03bb4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure14-1.png",
        "caption": "Fig. 14: Tube-to-plate simulation",
        "texts": [
            " The distance from the end of tool to the seam is set to 10 mm, and the welding speed is set to 10 mm/s. The plates and the welding trace of the master robot are shown in Fig. 10. Simulation results are drawn in Fig. 11 to 13. In Fig. 11 the endpoint trajectories of the master and slave robot in world coordinate system is shown, and joint angels of the master and slave robot are shown in Fig. 12 and Fig. 13 respectively. All simulation results can show the effectiveness of our proposed method. The model parameters of tube-to-plate and welding discrete seam are shown in Fig. 14. Simulation results are drawn as follows. Joint angels of the master and slave robot are shown in Fig. 15 and Fig. 16 respectively. All simulation results can show the effectiveness of our proposed method. In this simulation, the model parameters of tube-to-tube and the workpiece coordinate system are shown in Fig. 17(a). And the workpiece coordinate system, the master robot tool coordinate system and the master robot flange coordinate system are coincident with each other. In addition, the workpiece coordinate system is fixed to and moves along tube-to-tube plate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001803_978-3-319-07944-8_18-Figure18.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001803_978-3-319-07944-8_18-Figure18.2-1.png",
        "caption": "Fig. 18.2 3D schematic diagram of fi lament winding process",
        "texts": [
            " Peters and Humphrey ( 1987 ) reported that the fi lament winding method was fi rst presented in the form of patent since 1963. In 1964, a monograph describing the fi lament winding process was published. An automated fi lament placement technology was introduced in 1990 (Anon 2011 ). The fi lament winding process involves drawing continuous fi bres through a container of resin mixture or bath and winding the continuous resin impregnated fi bres around a rotating mandrel to achieve the desired shape (Fig. 18.2 ). The fi bres are placed on the rotating mandrel by means of a horizontal carrier. The fi bre orientation is controlled by adjusting the speed of the carrier. Mandrel can be removable, can be sacrifi cial or can form as part of the component. Thereafter, the component is cured in an oven for a certain period of time at a suitable temperature. Defects are normally found in products made from fi lament winding process such as in the forms of voids, fi bre wrinkles and delaminations (Mallick 2008 )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002571_2168-9806.1000135-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002571_2168-9806.1000135-Figure3-1.png",
        "caption": "Figure 3: Dynamic equilibrium of the rigid crawler shoe, i.",
        "texts": [
            " Governing equations of motion and solution methodology The kinematic and dynamic equations of motion that govern the propelling motion of the rigid crawler track on oil sand terrain based on multi-body dynamics theory [15,16] can be found in Frimpong and Thiruvengadam [12,13]. Frimpong and Thiruvengadam [12,13] presented kinematic equations governing pin joints between adjacent crawler shoes, spherical joints between adjacent oil sand units and motion constraints applied on crawler shoes and oil sand units. The free body diagram of a crawler shoe i with inertia forces in dynamic equilibrium with external and joint constraint forces is shown in Figure 3 [16,17]. The external forces acting on the crawler shoe # i are the gravity force (mig) due to self-weight of the shoe, uniformly distributed load (wi) due to machine weight and contact forces ,i i c cF T due to interaction between crawler shoe i and ground as shown in Figure 3. The joint forces at the spherical joints ( 1, 1,,\u2212 \u2212i i i i s sF T and , 1 , 1,+ +i i i i s sF T ) and parallel primitive joints ( 1, 1,,\u2212 \u2212i i i i p pF T and , 1 , 1,+ +i i i i p pF T ) are also shown in Figure 3 For 63 interconnected rigid multi-body system shown in Figure 2, the differential-algebraic equations of motion can be written from Shabana and MSC [16,18] as in equation (1). + = + T u a gMu K Q Q\u03bb ( , ) 0=tK u (1) M - Mass matrix of the system; u - vector of system generalized coordinates; K - constraint equations due to joints and applied motion; Qa - Applied forces; Qg - gyroscopic terms of the inertia forces. Frimpong and Thiruvengadam [13] presented detailed formulation for the generalized inertia and external forces acting on the crawler shoes and oil sand unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002838_adv.2020.125-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002838_adv.2020.125-Figure3-1.png",
        "caption": "Figure 3: Schematic of 192 dogbone orientations in the build tray. (left) isometric view; (top-right) profile view; and (bottom-right) rear view. The vertical base-to-top pillar in each view indicates where EOS temperature controls scan the powder bed and indicates the front/door side of the build tray.",
        "texts": [
            "B Data Collection The printer used was an EOS P110 using RP Tools Slicing and PSW for communication. PA2200 polyamide powder was used at a 50:50 fresh to single-run powder ratio as per manufacturer specifications. The PSW settings were set at the PA2200-Balance settings with a 120m layer thickness with the build chamber nominal temperature set to 168C whereas the layer thickness was fixed at 100m layer thickness within RPW. A total of 192 Type 1BA ISO 527 standard dogbones were printed, four sets of 48 in positions labeled horizontal, vertical, angular, and horizontal-vertical (Figure 3). ht tp s: // do i.o rg /1 0. 15 57 /a dv .2 02 0. 12 5 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. C ol um bi a U ni ve rs ity L ib ra ri es , o n 10 M ar 2 02 0 at 2 1: 42 :0 5, 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/ Imaging of polyamide-12 (nylon) specimens was performed by X-ray computed tomography (XCT) [13-15] on an X-Tek/Metris XTH 320/225 KV scanner Nikon e ology Belmon CA . XCT a a was collec e a 80kV an 350\u03bcA X-ray power, each consisting of 3000 projections collected with 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002434_s11426-016-0243-7-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002434_s11426-016-0243-7-Figure2-1.png",
        "caption": "Figure 2 Schematic illustration of the interaction of the PAA/PAH films with the PDDA-modified substrates after being immersed in water with pH of <4.5 (a), >6 (b) and 5.5 (c) (color online).",
        "texts": [
            " The formation of the wrinkles in a swollen film is determined by the competing forces of the compressive stress () and the substrate confinement (s). The ionization ratio of carboxylate groups in PAA is lower than 50% in pH 4.5 water [37]. When the cross-linked (PAA/PAH)*n films are immersed into a more acidic water solution (pH<4.5), the interaction between the bottom PAA layer of the (PAA/PAH)*n films and the PDDA-modified substrates is largely weakened. Upon water adsorption, the compressive stress generated by film swelling exceeds the substrate confinement stress, leading to a complete exfoliation of the films from substrates (Figure 2(a)). The ionization ratio of carboxylate groups in PAA is ca. 100% in pH 6.5 water [37]. The electrostatic interaction between the bottom PAA layer of (PAA/PAH)*n films and the PDDA-modified substrates cannot be weakened when the films are immersed in a more basic water solution (pH>6.5). In this case, the compressive stress generated by film swelling is smaller than the substrate confinement stress (<s), prohibiting the formation of wrinkles (Figure 2(b)). When the cross-linked (PAA/PAH)*n films are in water solutions with pH of ca. 5.5, the interaction of the films with the PDDA-modified substrates is moderately weakened to a threshold value that allows for partial separation of the film from the underlying substrate. Upon water adsorption, the separated regions of the film buckle, while the left regions still adhere to the underlying substrates. In this way, wrinkles are generated on the cross-linked (PAA/ PAH)*n films (Figure 2(c)). With the decrease of the internal stress during a film deswelling process, the (PAA/ PAH)*n film reattach to the substrate. In a control experiment, a bilayer of PSS/PAH film was pre-deposited onto a PDDA-modified substrates before the deposition of (PAA/ PAH)*30 films. After the wrinkle-inducing step of immersing the (PSS/PAH)*1/(PAA/PAH)*n film in pH 5.5 water solution, no wrinkles can be formed when the RH of the environment is increased to 80%. This is because that the ionization of PSS is not governed by pH of water solution, and the interaction of the (PSS/PAH)*1/(PAA/PAH)*n film with the PDDA-modified substrate cannot be weakened after immersion in pH 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002896_s0005117920020058-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002896_s0005117920020058-Figure1-1.png",
        "caption": "Fig. 1. Geometrical interpretation of Theorem 1.",
        "texts": [
            " Consider a geometrical interpretation of two cases depending on the shape of the function S(x) or S\u22121(x) in Theorem 1. Denote by F1 the flow of the vector field |grad{S(x)}|f(x) through the surface \u0393 = {x\u2208D : S(x) = C} with the unit normal vector 1 |grad{S(x)}|grad{S(x)} and by F2 the flow of the vector field |grad{S\u22121(x)}|f(x) through the surface \u0393inv = {x \u2208 D : S\u22121(x) = C} with the unit normal vector 1 |grad{S\u22121(x)}|grad{S\u22121(x)}. A geometrical interpretation of both cases for x \u2208 R 2 is presented in Fig. 1: the functions S(x) and S\u22121(x) (the left part of Figs. 1a and 1b) as well as the flows F1 and F2 of the vector fields |grad{S(x)}|f(x) and |grad{S\u22121(x)}|f(x), respectively, through the corresponding level surfaces \u0393 and \u0393inv (the right part of Figs. 1a and 1b). If system (1) is stable, then the vector field flow F1 (F2) through the surface \u0393 (\u0393inv, respectively) takes a negative value (a positive value, respectively). Proof. In accordance with [29, Theorem 4.17], if x = 0 is an asymptotically stable equilibrium of system (1), then there exists a continuously differentiable and positive definite function S(x) such that S(x) \u2192 \u221e as x \u2192 D\u0304, grad{S(x)}Tf(x) < 0 for any x \u2208 D \\ {0} and grad{S(x)}Tf(x) \u2223\u2223\u2223 x=0 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001836_iecon.2015.7392422-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001836_iecon.2015.7392422-Figure2-1.png",
        "caption": "Fig. 2. Concept of modulated potential field.",
        "texts": [
            " In Section II, the proposed modulated potential field is described. In Section III, the direct kinematics for position and force control are described. In Section IV, the simulation and experiment was conducted to confirm the validity of analysis results. Finally, 002166 978-1-4799-1762-4/15/$31.00 \u00a92015 IEEE this paper is concluded in Section V. In this paper, modulated potential field for position adjusting with human interaction is defined. The required performance of modulated potential field is shown in Fig. 2. There are three requirements that described as follow. 1) Far from the Destination: Physician can Free moving the manipulator. And the trajectory will be decided by human. 2) Close in the Destination: The manipulator will apply adjusting force for easy converge to the destination. 3) Over the Destination: The manipulator will apply adjusting force for collision avoidance. The Lennard-Jones potential (L-J potential) can be suitable with these three requirements. It is a mathematically simple model that approximates the interaction between a pair of neutral atoms or molecules"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000508_j.jsv.2015.07.002-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000508_j.jsv.2015.07.002-Figure2-1.png",
        "caption": "Fig. 2. Direction and contact force vectors at a contact point on the thread, (a) side view and (b) section a a'.",
        "texts": [
            " Then, the propensity of the mode-coupling type instability is investigated by the two-mode models of the torsion/axial mode pair and the bending mode pair without loss of generality. Particularly, it is also proven that the propensity of the mode-coupling instability in the torsion/axial mode pair can be estimated from only tooth topology. The rotation of the lead screw makes the nut move in the axial direction where the nut normally carries an weight as shown in Fig. 1. During operation, contact forces act on the contact thread between the screw and nut. Their direction vectors at a contact point are defined as in Fig. 2. The following notations are used for the description of the lead screw system: rotating speed (\u03a9), effective radius of screw (R), helix angle (\u03b2), half of thread angle (\u03b3), axial load \u00f0Fz\u00de, nut length (b) and its location \u00f0zo\u00de. In the model, the positivity of \u03a9 determines the direction of rotation in such a way that a positive value of \u03a9 represents the forward rotation in Fig. 3. It is assumed that the contact stiffness per unit area kc\u00f0 \u00de is uniformly distributed over the contact thread as illustrated in Fig. 3. Contact stress acting on the thread is also assumed to be uniformly distributed over the contact side of the section a a' (Fig. 2b) so that the contact kinematics at c (Fig. 3) is used for the description of the contact stress variation on the contact thread. For the bending mode, the Bernoulli\u2013Euler beam theory is applied in the analysis [22]. Please cite this article as: J. Kang, Friction-induced noise of gear system with lead screw and nut: Modecoupling instability, Journal of Sound and Vibration (2015), http://dx.doi.org/10.1016/j.jsv.2015.07.002i Fig. 3. Description of contact location and stiffness, (a) contact location \u2018c\u2019 and (b) contact stiffness per area, kc ",
            " It reduces the contact displacement and contact normal stress variation as uc \u00bcw\u00f0z; t\u00dek\u00feR\u03be\u00f0z; t\u00dee\u03b8: (26) ~N \u00bc kc R cos \u03b3 sin \u03b2\u00f0 \u00de\u03be\u00fe cos \u03b3 cos \u03b2\u00f0 \u00dew : (27) Besides, two mode discretization is composed of one axial and torsion mode as w\u00f0z; t\u00de \u00bc\u03a6z m\u00f0z\u00deUq1\u00f0t\u00de; (28) \u03be\u00f0z; t\u00de \u00bc\u03a6\u03b8 n\u00f0z\u00deUq2\u00f0t\u00de; (29) The equations of motion for two-mode discretization take the form of \u20acq1 \u20acq2 ( ) \u00fe \u03bc \u03a9j j d1 d2 d2 d3 \" # _q1 _q2 ( ) \u00fe \u03c92 a 0 0 \u03c92 t \" # \u00fekc a1 a2 a2 a3 \" # \u00fe\u03bckc b1 b2 b3 b4 \" # ! q1 q2 ( ) \u00bc 0; (30) where \u03c92 a and \u03c92 t stand for the natural frequencies of the uncoupled axial and torsion modes without contact, respectively, and other components of the matrices are described in (A.1)\u2013(A.10) of Appendix A. It is highlighted that the damping elements (A.1)\u2013(A.3) of Appendix A in the damping matrix come from friction stress variation perpendicular to the steadysliding direction es in Fig. 2. Since the displacement perpendicular to es is composed of the radial, axial and circumferential components, the friction damping elements in (A.1)\u2013(A.3) depends on the axial and circumferential displacements of the axial and torsion modes, respectively. Here the radial component does not show up due to the lack of the radial displacement in the two-mode model including the torsion and axial modes. In contrast, the radial and axial components will induce the friction damping in the bending mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003332_aim43001.2020.9158809-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003332_aim43001.2020.9158809-Figure5-1.png",
        "caption": "Figure 5. Dynamic obstacle velocity potential field",
        "texts": [
            " The total repulsion velocity from the thi robot\u2019s neighbor is defined as: ( , , ) i ji i i k Ok L L RR RR j N rad\u2208 = \u2211 r o v v (12) D. Collision Avoidance between Obstacles When facing a static obstacle, we will use the velocity field of a circular equipotential line similar to the previous part. When facing a dynamic obstacle, we will stretch the original circular equipotential line in the direction of the obstacle velocity. The equipotential lines of the velocity field are trough-shaped, which is shown in Figure 5. The velocity potential function of the thk obstacle ( )kH o is defined as: 1( ) ln( ) 2 2 kcollision k k collision f radH radd \u03b1     = \u2212 \u2212   \u2212   o o o 1437 Authorized licensed use limited to: Cornell University Library. Downloaded on August 30,2020 at 03:02:18 UTC from IEEE Xplore. Restrictions apply. 2 2 if < 2 k L o k collision k Od rad d\u03b1   +  \u2212  <     o o o\uff0c (13) The repulsion velocity of the thk obstacle to the thi robot is defined as the negative gradient of ( )kH o : = ( ) ki L RO kH\u2212 \u2207v o (14) 1 2 1 1 1 1[( ) ( )] 2 2 + ( ) if < 0 if ki L k L L collision collisionOk k L L RO L Lk O k collision k OL k L k O rad radd d rad d d \u03b1 \u03b1  \u2212 \u2212 \u2212  =  \u2212 <   \u2265 o o o v o o o o o \uff0c \uff0c (15) where Od is the safe distance from the obstacles, L ko is the position vector of the thk obstacle in the local coordinates. The coefficient 1 2, 0\u03b1 \u03b1 > play the same role in Eq. (13) and Eq. (15) as it does in Eq. (9) and Eq. (11). For the dynamic obstacle, we extend the velocity potential field in the direction of kOv , The area around the obstacle is divided into three areas in Figure 5. In each area, L ko in Eq. (15) is calculated by: , , , k i i L k i k k k i k i k i i O R R O R O MO R O M R O M MR R  \u2208\u0399  \u22c5= \u2212 \u2208\u0399\u0399   \u2208\u0399\u0399\u0399 o        (16) The stretch vector kO M  is defined as follows: 2 O k OO M t \u03b2 = \u22c5\u2206 \u22c5 v v  (17) where 0\u03b2 > is a coefficient to adjust the range of the dynamic obstacle and t\u2206 is the interval time. As the obstacle speed Ov increases, the stretch vector kO M  increases at the same time. This means that the range of dynamic obstacles will increase. The velocity potential field staticH and dynamicH is shown in figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure1-1.png",
        "caption": "Figure 1. The circular toroid obtained as a surface trajectory in a RR dyad.",
        "texts": [
            " The four circular generatrices of an OCT are associated with four RRS open chains producing the same Cartesian equation of an OCT. Some assemblies of two RRS open chains yield a geometrical derivation of four types of Bennett linkages having the same link lengths and the same absolute value of the link twist angles. An appropriate convention for orientating vectors and angles is adopted to discuss special cases comprehensively as well. Let us consider the serial concatenation of two revolute R pairs shown in Figure 1, which is sometimes called also a two-hinge dyad in the literature on mechanism theory. The open RR chain includes three rigid bodies connected in series by two revolute R joints or hinges. Each of the two R joints (or R pairs) is characterized by its axis and each axis is determined by the datum of any one of its points and a unit vector parallel to the axis in an initial home configuration of the RR chain. Assuming that the two axes are not parallel, there is a unique common perpendicular to the two axes",
            " It is convenient to choose two points on the common perpendicular to specify the two R axes. When the first fixed R pair is locked, a point P belonging to the end body of a two-hinge dyad traces a circle on a plane, which is perpendicular to the second R axis and is located at a distance d of the common perpendicular. The distance d is called offset. When the fixed R is unlocked, the point P traces a surface of revolution, which is called at University of Birmingham on April 21, 2015pic.sagepub.comDownloaded from a general circular toroid, as shown in Figure 1. In the noteworthy paper, Fichter and Hunt13 proved that a plane that is bitangent to a general circular toroid intersects the surface along a couple of congruent circles. The pertinent proof is established by using the complex projective completion of the Euclidean space. In what follows, one will not consider imaginary geometric entities but will focus on the special case with a zero offset. When the offset (d) is zero and two R axes are not twisted by a right angle (j j 6\u00bc 90 ), the circular toroid becomes an OCT depicted in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure5.10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure5.10-1.png",
        "caption": "Figure 5.10 Slack loop system schematic, vacuum column",
        "texts": [
            " Servoing the high inertia \u201csupply\u201d roll to meet these two requirements simultaneously can be difficult. A direct approach is to separate the requirements as shown schematically by the examples in Figures 5.10 and 5.11. 268 Electromechanical Motion Systems: Design and Simulation R M CONTROL VELOCITY COMMAND \u03b8 Figure 5.11 Slack loop system schematic, dancer arm A \u201cslack loop\u201d of the material is created between the supply reel and the input/output mechanism in order to provide time for the supply reel to accelerate to the input/output velocity. In Figure 5.10 the slack loop is formed in an evacuated column (a vacuum column); in Figure 5.11 the slack loop is formed by guiding the material around a roller at the end of a spring loaded lever arm. The position of the loop is divided into three regions, delineated by two position sensors, as CW Drive, Brake and CCW Drive. Depending on the loop position, the supply motor is powered only by maximum CW or CCW voltage, namely the term \u201cBang-Bang\u201d. An example from second generation digital tape handler equipment demonstrates the operation of such a \u201cBang-Bang\u201d system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000792_jproc.2014.2355714-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000792_jproc.2014.2355714-Figure7-1.png",
        "caption": "Fig. 7. Electrical equivalent circuit model of Fick\u2019s first law of diffusion.",
        "texts": [
            " Fick\u2019s first law of steady state diffusion defines diffusion flux \u00f0J\u00de in terms of the diffusivity parameter \u00f0D\u00de and the concentration \u00f0c\u00de gradient in the steady state [67]. In 1-D, Fick\u2019s first law is given by J \u00bc D @c @x : (3) Multiplying by the cross section of the diffusion channel A, Fick\u2019s first law can be restated in terms of a diffusion current ID \u00bc A J \u00bc A D @c @x : (4) Assuming a uniform steady state concentration gradient in the diffusion channel, an equivalent lumped electrical circuit model based on charge drift and Ohm\u2019s law can be developed to satisfy Fick\u2019s first law, as given in Fig. 7. The concentrations are modeled with independent sources and the diffusivity with a diffusion resistance. Integration of (4) under steady state gives the resistance in RD \u00bc c ID \u00bc L A 1 D : (5) This diffusion resistance has the unit of s/m3 and is the mathematical inverse to an effective volumetric flow rate of fuel to the electrode. Additionally, the amount of fuel in the diffusion channel must be considered. In the equilibrium case of Fig. 7 with c0 equal to c1, the concentration will be constant throughout the diffusion channel and the total number of moles of fuel stored in the diffusion channel is given by the product of volume of the diffusion channel and the concentration. To transition to an electrical simulation model of Fick\u2019s laws, conversion Table 2 is established to relate diffusion model and electrical circuit model quantities. Fick\u2019s first law model will only be valid under steady state and slowly varying conditions. It does not capture the diffusion limited dynamics of interest, and Fick\u2019s second law needs to be considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002959_aeat-05-2019-0094-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002959_aeat-05-2019-0094-Figure4-1.png",
        "caption": "Figure 4 Oil-film bearing balance position",
        "texts": [
            " The kinetic and strain energy of the shaft unit are expressed as follows: Tr \u00bc 1 2 \u00f0 l 0 rA _V r 2 1 _Wr 2 1 _Ur 2h in o ds1 \u00f0 l 0 rIrD _Br 2 1 _Cr 2h in o ds 1 2 \u00f0 l 0 X1 _ar\u00f0 \u00derIrP _CrBr _BrCr ds1 1 2 \u00f0 l 0 X1 _ar\u00f0 \u00de2rIrPds (7) Double-helical geared rotor system Ying-Chung Chen, Tsung-Hsien Yang and Siu-Tong Choi Aircraft Engineering and Aerospace Technology Volume 92 \u00b7Number 4 \u00b7 2020 \u00b7 653\u2013662 where A is the cross-section area, r is the mass density of the shaft, IrD and IrP are the transverse moment of inertia and polar moment of inertia of the shaft, respectively, E is Young\u2019s modulus of the shaft, k is shear factor for the circular crosssection of the shaft andG is the shearmodulus of the shaft. The equation of motion of the finite shaft element can be derived by substituting kinetic energy (7) and strain energy (8) into Lagrange\u2019s equation: Mr\u00bd f\u20acqrg1X Gr\u00bd f _qrg1 Kr\u00bd fqrg \u00bc f0g (9) where [Mr], [Gr] and [Kr] are the mass, gyroscopic effect and stiffness matrices, respectively, of the finite shaft element. Oil-film bearing In Figure 4, e denotes the eccentricity of the bearing, C the radial clearance of the bearing and F the radial load of the bearing. In this paper, only half of the bearing was assumed to be covered by oil film, which signified that it did not cover the entire bearing duringmovement. The stiffnessKij and damping Cij coefficients of the bearing are as follows: Kij \u00bc Wkij C ; Cij \u00bc Wcij XC (10) In (10), Kij is the dimensionless stiffness coefficient and Cij the dimensionless damping coefficient. Kij and Cij are expressed as follows: k11 \u00bc 4 p2 1 321p2\u00f0 \u00de\u00ab2 1 2 16 p2\u00f0 \u00de\u00ab4 Q \u00ab\u00f0 \u00de 1 \u00ab2\u00f0 \u00de ; k12 \u00bc p p2 1 321p2\u00f0 \u00de\u00ab2 12 16 p2\u00f0 \u00de\u00ab4 Q \u00ab\u00f0 \u00de \u00ab ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 \u00ab2\u00f0 \u00dep k21 \u00bc p p2 2p2\u00ab2 16 p2\u00f0 \u00de\u00ab4 Q \u00ab\u00f0 \u00de \u00ab ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 \u00ab2\u00f0 \u00dep ; k22 \u00bc 4 2p2 1 16 p2\u00f0 \u00de\u00ab2 Q \u00ab\u00f0 \u00de (11) c11 \u00bc 2p p2 12 24 p2\u00f0 \u00de\u00ab2 1p2\u00ab4 Q \u00ab\u00f0 \u00de \u00ab ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 \u00ab2\u00f0 \u00dep ; c12 \u00bc 8 p2 1 2 p2 8\u00f0 \u00de\u00ab2 Q \u00ab\u00f0 \u00de c21 \u00bc 8 p2 1 2 p2 8\u00f0 \u00de\u00ab2 Q \u00ab\u00f0 \u00de; c22 \u00bc 2p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 \u00ab2\u00f0 \u00dep p2 12 p2 8\u00f0 \u00de\u00ab2 Q \u00ab\u00f0 \u00de \u00ab (12) Q \u00ab\u00f0 \u00de \u00bc p2 1 \u00ab2\u00f0 \u00de116\u00ab2 3=2 ; \u00ab \u00bc e C (13) where \u00ab is the eccentricity ratio of the bearing obtained from the bearing characteristic coefficient S using the following formula: S \u00bc mX W R C 2 Ld 3 D \u00bc 1 \u00ab2\u00f0 \u00de2 p\u00ab ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi p2 1 \u00ab2\u00f0 \u00de1 16\u00ab2 p (14) where Ld is the length of the bearing, R is radius of the shaft, D is diameter of the shaft, m is lubricant viscosity of the bearing and X is spin speed of the shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000518_eeeic.2015.7165426-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000518_eeeic.2015.7165426-Figure11-1.png",
        "caption": "Fig. 11. Flux lines and flux density map at different time moments",
        "texts": [
            " In terms of visualization and presentation of results, FEA is by far on the first place. Of course, FEA results, based on the accuracy of the achieved numerical model, can be considered only after the validation of the model. Therefore it is necessary first to compare some FEA results with experimental ones. Once a numerical model and analysis technique are validated, FEA allows a multitude of information on the studied machine, even if the shape or the materials of some components are changed. For example, in Fig. 11, the magnetic field lines and magnetic flux density map of the studied generator, at different time moments, are presented. In the first moment when shortcircuit occurs, the stator winding current increases suddenly. Because the field winding current also increases suddenly, the magnetic flux density remains relatively constant, Fig. 11c. Only after the transient regime is finished, when the generator operates in steady state short-circuit, the magnetic flux density changes essentially, showing that the generator is magnetically reduced loaded, Fig. 11d. The transient torque (for half of generator) in the three cases is showed in Fig. 12. The torque, as the currents, is maximum for single-phase short-circuit and then follows twophase and three-phase short-circuits. [1] C Bassi, D. Giulivo, A. Tessarolo, \u201cTransient finite-element analysis and testing of a salient-pole synchronous generator with different damper winding design solutions,\u201d XIX International Conference on Electrical Machines, ICEM , Rome, 2010. [2] Rotating electrical machines, part 4: Methods for determining synchronous machine quantities from tests, IEC 34-4 Std, 1995"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001042_2014-01-1664-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001042_2014-01-1664-Figure1-1.png",
        "caption": "Figure 1. Rotary engine sealing system",
        "texts": [
            "he rotary engine provides a high power density and low vibration when compared to the piston engine. This makes the rotary engine a prime candidate for sports cars due to its small weight and for hybrid vehicles where packaging is critical. However, its trochoidal shape leads to a more complex system to seal the gas mixture in the different chambers and crankcase oil, as shown in Figure 1. Additionally to the oil injected to lubricate the apex, side and corner seals, oil can leak from the crankcase to the combustion chamber through the inner and outer oil seals, defined here as internal oil combustion (IOC). IOC accounts for about half the total oil consumption, which is far greater than the necessary amount of oil to lubricate the contact between the oil seals and the side housing. A better understanding of the oil transport mechanisms through the oil seals would help reduce IOC, as the oil seals are the main barrier between crankcase oil and the combustion chamber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002183_2168-9792.1000163-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002183_2168-9792.1000163-Figure5-1.png",
        "caption": "Figure 5: Electronic components connection.",
        "texts": [
            " An RF transmitter receives serial data and transmits it wirelessly through RF through its antenna connected at pin 4. The transmission occurs at the rate of 1Kbps - 10Kbps. The transmitted data is received by an RF receiver operating at the same frequency as that of the transmitter. The RF module is used with a pair of encoder or decoder. The encoder is used for encoding parallel data for transmission feed while reception is decoded by a decoder. The assembly and connections of the various electronic components of the Quadcopter is shown in Figure 5. The fabricated quadcopter model is shown in the Figure 6. The total mass of the quadcopter is estimated in Table 2. The total empty mass estimated from the above table is about 901 grams. As the expected payload capacity is considered as 300 grams, the quadcopter should be able to fly with a total mass of around 1200 grams. [10-13] The thrust of the quadcopter [4] is given by the equation T = \u03c0D2\u03c1v\u0394v/4 Where T is thrust in N, D is Propeller diameter in m, \u03c1 is Density of the air \u2013 1.22 kg/m3 Also V = \u0394V/2 Where V is the velocity of air at the propeller, \u0394V is the velocity of the air accelerated by propeller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001532_0954406214536700-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001532_0954406214536700-Figure8-1.png",
        "caption": "Figure 8. Three-order elliptic helical gear shaping by the optimal linkage-model.",
        "texts": [
            "comDownloaded from Gear shaping experiment and accuracy analysis According to strategy (4) and linkage-model (23), a shaping module for external non-circular helical gears has been developed based on a shaping platform with advanced RISC machines (ARM) and digital signal processor (DSP) and field programmable gata array (FPGA).24 The shaping module is implanted in a NC system and has been used in a gear shaper (STAR2010C). A three-order elliptic helical gear aforementioned is shaped, as shown in Figure 8. The processing parameters are consistent with those in simulation. The result shows that the gear can be shaped correctly. As shown in Figure 3, as for strategies (1), (2) and strategies (5), (6), vt becomes maximum while equals to p/3, p, 5p/3, respectively, yet minimum while equals to 0, 2p/3, 4p/3. As for strategies (7), (8), vt becomes maximum while roughly equals to 0, 2p/3, 4p/3. respectively, but minimum while roughly equals to p/3, p, 5p/3. In order to verify the consistency of the accuracies of the tooth surface shaped by strategy (4) and linkage-model (23), the surface roughness of the 1st tooth ( \u00bc 0) and the 7th tooth ( \u00bc p/3) of the gear shown in Figure 8 are detected by a surface roughness measuring instrument.25 The point cloud of roughness for the two tooth surfaces is shown in Figure 9. It is obvious that the roughness of the 1st tooth is very close to the 7th one. It can be concluded from the result that all tooth surface have similar roughness. Hence, all of the accuracies of the tooth surface in the gear are uniform. This test has verified the simulation result. That is to say, cutting marks among every tooth shaped by the optimal linkage-model are well-distributed, and then the accuracy among every tooth surface is uniform. Form errors of tooth profiles Tooth profile error is also an important factor influencing the transmission accuracy of the gear pair. In order to verify the consistency of the accuracies of the tooth profile in the three-order elliptic helical gear shaped by strategy (4) and linkage-model (23) in Figure 8, the profile forms of the left tooth surfaces of the 1st tooth, the 4th tooth and the 7th tooth of the gear are detected by a profile measuring instrument.26 Moreover, those of the right tooth surfaces are detected in the same way. Compared the measured results to the theoretical model, all the form errors of tooth profiles can be obtained, as shown in Figure 10. The profile errors of the three left tooth surfaces are 11.9mm, 10.8 mm, 10.7 mm, respectively. The average of them is 11.1mm. Those of the three right tooth surfaces are 11"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002589_ichve.2016.7800782-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002589_ichve.2016.7800782-Figure3-1.png",
        "caption": "Fig. 3. phase-ground air gaps",
        "texts": [
            " Test sample photo is as shown in Figure 2, including upper 2-phase cross arm and 1-phase cross arm in middle layer. Four kinds of grounding methods for ground wire were designed aiming at getting insulation feature of composite tower and poles. 978-1-5090-0496-6/16/$31.00 \u00a92016 IEEE 1 Grounding method 1 Grounding method 1 was that ground wire was led to ground along pole top, and phase to ground gap of tower and pole head was wire (clamp)-loop. Schematic diagram of grounding type and phase to phase gap were as shown in Figure 2 and Figure 3 respectively. 2) Grounding method 2 Grounding method 2 was that ground wire was grounded by leading wire down from pole top and after passing through series gap. Gap distance between the two electrode of the series gap was 0.75 m. The distance of two 2 fixed points in the pole body was 1.0 m as shown in Figure 4. 3) Grounding method 3 Grounding method 3 was that ground wire was routed away from pole body along wire direction from pole top, and led down to ground as shown in Figure 5. 4) Grounding method 4 Grounding method 4 was that ground line was routed away from pole body in vertical wire direction from pole top, and led down to ground in the outside of wire as shown in Figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003488_icra40945.2020.9196838-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003488_icra40945.2020.9196838-Figure1-1.png",
        "caption": "Fig. 1: McFoamy agile fixed-wing platform",
        "texts": [
            " Fixed-wing UAVs, which generate lift through their wing and other lifting surfaces, are the traditional choice for tasks that require large coverage, endurance, and high-speed flight. However, compared to rotorcraft, their maneuverability is limited: they are incapable of hovering in place, and often require additional infrastructure to take-off and land. A class of fixed-wing aircraft, equipped with propulsion systems with a large thrust-to-weight ratio, coupled with high aspect ratio wings and large control surfaces that capitalize on propeller slipstream, like the one in Fig. 1, have demonstrated the capacity to perform aggressive, precise aerobatic maneuvers [1]. These platforms, known as agile fixed-wing aircraft, bridge the gap between conventional fixed-wing aircraft and rotorcraft. Their unique design allows them to remain airborne even when the lifting surfaces no longer generate sufficient lift, while retaining control authority through propeller slipstream over the control surfaces. Indeed, the slipstream greatly enhances the motions of agile UAVs as demonstrated in [2]",
            " It allowed this process to take place in a safe laboratory environment, rather than risking damage to the physical platform at this precarious stage of development. Following this, experimental tests were undertaken to further adjust the control gains, as needed. The position control gains required minimal adjustment from those in the SITL, while the attitude control gains had to be reduced to compensate for the effects of sensor noise. The final values of the experimental gains, kept constant throughout all the field trial maneuvers discussed next, are given in Table I. An off-the-shelf WM Parkflyers McFoamy RC aircraft, shown in Fig. 1, with mass of 0.45 kg and 0.86 m wingspan was employed for the field trials. Other aircraft properties can be found in [8], [7]. The control algorithm was implemented within the Pixracer embedded autopilot hardware. This system uses a 180 MHz ARM Cortex M4 processor and runs Px4 software allowing for attitude and angular velocity estimation with built-in sensors, and position and velocity estimates when augmented with a GPS unit. The experiments were conducted outdoors with a mean wind of about 15 km/h at an angle between 100\u25e6 and 140\u25e6 from true north, with gusts up to 30 km/h"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001572_s11071-013-1014-5-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001572_s11071-013-1014-5-Figure6-1.png",
        "caption": "Fig. 6 The coordinate of drilling shaft in deep hole drilling",
        "texts": [
            " [6] investigated the influence of vortex motion of spinning drilling shaft caused by the hydro-force of cutting fluid. Kong et al. [7] established the concentration mass dynamic model of drilling shaft system with the multi-degrees, and proposed a method to calculate the nonlinear hydrodynamic forces of cutting fluid. Simultaneously, this model is feasible reported by the experimental results. Based on a similar idea, the coordinate of drilling shaft in MQL deep hole drilling for calculation is shown in Fig. 6 and the drilling shaft equations of lateral motion can be represented as follows: \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 mx\u0308 + 768EI 7L3 x = Fx + Fcx sin\u03c9t + mex\u03c9 2 cos\u03c9t + mey\u03c9 2 sin\u03c9t my\u0308 + 768EI 7L3 y = Fy + Fcy cos\u03c9t + mg + mey\u03c9 2 cos\u03c9t \u2212 mex\u03c9 2 sin\u03c9t (17) In (17), m is the lumped mass of the drilling shaft; E and I are elastic modulus and moment of inertia, respectively; L is the length of the drilling shaft; Fx and Fy are nonlinear hydrodynamic forces of MQL cutting fluid in x and y directions, respectively, Fx =\u222b\u222b \u03a9 p cos\u03d5 d\u03a9 and Fy = \u222b\u222b \u03a9 p sin\u03d5 d\u03a9;\u03a9 is the cutting fluid field in the drilling hole; p is the pressure distribution of MQL fluid in \u03a9 and it can be obtained from (16), which is needed at every step of solution process when the drilling shaft trajectory of (17) is solved; g is the acceleration of gravity; ex and ey are the transverse displacement of the drilling shaft in x and y directions, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure3-1.png",
        "caption": "Fig. 3: Welding torch coordinate system",
        "texts": [
            " The attitude of the welding seam at the current discrete point is defined as follows. Weld slope S is the angle between the plane XWOWYW of {FW} and axis XP, and its value ranges from -180 to 180 degree. Weld rotation R is the angle which makes the plane XPOPYP of {FP} parallel to axis ZW of {FW} through rotating {FP} around axis XP, and its value ranges from -180 to 180 degree. Weld deflection D is the angle between the plane XWOWZW of {FW} and axis XP of {FP}, and its value ranges from -180 to 180 degree. The welding torch coordinate system {FT} is shown in Fig. 3. The origin OT is chosen as the endpoint of the welding torch; axis XT coincides with the axis of welding torch and points to the welding seam; axis ZT lies in the plane determined by axis XT and the centerline of the flange, and perpendiculars to axis XT but deviations to the flange centerline; axis YT is determined by the right-hand rule. In order to describe the pose of the welding torch more clearly, the welding angle, the moving angle and the rotation angle are defined as Fig. 4 shows. The welding angle W is the angle between axis XT of {FT} and axis YP of {FP}, and the direction is from YP to -XT"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure8-1.png",
        "caption": "Fig. 8 Kinematic Sketch of the Segment",
        "texts": [
            " For the convenience of calculation, the coordinate of the segment k is set as the D-H coordinates numbered 2k by default. And then r j k , r j k\u22121 can be determined as: r j k = R2k \u03c1 j k + P2k, r j k\u22121 = R2(k\u22121) \u03c1 j k\u22121 + P2(k\u22121) (5) Where, \u03c1 j k is the position vector of the hole j on the lower plate of segment k in the D-H coordinate C2k . \u03c1 j k\u22121 is the position vector of the hole j on the upper plate of segment k \u2212 1 in the D-H coordinate C2(k\u22121). \u03c1 j k and \u03c1 j k\u22121 are constant vectors. The length of rope j is determined by d j k and \u03b7 j k in Fig. 8. Referring to Eqs. 2\u20135, the length of rope j can lj = \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 rem(j,5)\u2211 k=1 \u2225\u2225\u2225d j k \u2225\u2225\u2225 + l j 0 j = 1, 1 + K, 1 + 2K rem(j,5)\u2211 k=1 \u2225\u2225\u2225d j k \u2225\u2225\u2225 + rem(j,5)\u2211 k=1 \u2225\u2225\u2225\u03b7 j k\u22121 \u2225\u2225\u2225 + l j 0 j = 2, 3 \u00b7 \u00b7 \u00b7K; K + 2, 8 \u00b7 \u00b7 \u00b7 2K; 2K + 2, 13 \u00b7 \u00b7 \u00b7 3K (6) be calculated, as shown in Eq. 6. The procedure b is executed according to Eq. 2 and the procedure d is executed according to Eq. 6. Therefore, the procedure a is the inverse calculation of Eq. 6 and the procedure c is the inverse calculation of Eq. 2. Due to I = 10 > 6 in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002387_cae.21769-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002387_cae.21769-Figure4-1.png",
        "caption": "Figure 4 Quadcopter motion diagrams for (a) rolling, (b) pitching, (c) yawing.[Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com].",
        "texts": [
            " u \u00bc X4 i\u00bc1 f i \u00bc k X4 i\u00bc1 vi 2 \u00f08\u00de Multiplying Equation (1) by vector ~u \u00bc 0 0 u\u00bd T gives the total thrust ~FR expressed with respect to the inertial reference frame as shown in Equation (9). ~FR \u00bc sin w sin c\u00fe cos w sin u cos c cos w sin u sin c sin w cos c cos w cos u 2 64 3 75u \u00f09\u00de Substituting Equations (7), (9), and the weight vector ~w \u00bc 0 0 mg\u00bd T in (4) and rearranging give the nonlinear dynamic Equations (10)\u2013(12) for translational motion. \u20acx \u00bc sin w sin c\u00fe cos w sin u cos c m u kd _x 2 \u00f010\u00de \u20acy \u00bc cos w sin u sin c sin w cos c m u kd _y 2 \u00f011\u00de \u20acz \u00bc cos w cos u m u g kd _z 2 \u00f012\u00de Roll w, Pitch u, and Yaw c Torques. Looking at Figure 4a, a roll rotation is achieved by creating a speed difference Dv between motors 2 and 4 while the speeds of motors 1 and 3 remain the same. If the quadcopter has a nonzero roll angle, it translates laterally; either to the right or to the left. The torque difference between the left and right motors gives the rolling torque tw in Equation (13) where ti is the torque about the center of mass generated by the ith motor and l is the shortest distance in meters from the center of mass to any rotor\u2019s shaft. tw \u00bc t2 t4 \u00bc f 2 f 4\u00f0 \u00del \u00bc lk v2 2 v4 2 \u00f013\u00de A similar argument is applicable to pitch rotations (see Fig. 4b). It is achieved by creating a speed differenceDv between motors 1 and 3while the speeds ofmotors 2 and 4 remain the same. Longitudinal translation (forward or backward translation) is associated with any nonzero pitch rotation. Figure 5 depicts this motion. The pitching torque tu can be defined by tu \u00bc t3 t1 \u00bc f 3 f 1\u00f0 \u00del \u00bc lk v3 2 v1 2 \u00f014\u00de To make yaw rotations (see Fig. 4c), the torque about the zn axis must be unbalanced. To turn the quadcopter counterclockwise, the speed of the right-rotating motors (motors 2 and 4) is increased byDv and the speed of the left-rotating motors (motors 1 and 3) is decreased by the same speed differenceDv. The speed difference Dv must be equal to achieve pure turning about the local vertical axis zb. Equation (15) gives the yaw torque tcwhere tMi is the torque produced by the ith motor about its shaft, Jp is the moment of inertia of each propeller, and c is a constant known as force-tomoment scaling factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002135_1464419315618862-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002135_1464419315618862-Figure7-1.png",
        "caption": "Figure 7. Eleven DOFs vibration model.",
        "texts": [
            "16 In order to study the car body vibrations transferred by the mounting system, an elastic foundation comprising the car body, suspensions, wheels and tyres should be considered. Without regard to the lateral and roll movements of car body, the front suspensions, wheels and tyres are equivalent to a suspension, a wheel, and a tyre, respectively. And the same equivalent is processed for the rear suspensions, wheels and tyres. The longitudinal stiffnesses and dampings of the tyres and suspension bushings are equivalent to a spring-damper component located between the car body and the ground, and the equivalent stiffness and damping are denoted as kTL and cTL in Figure 7. So far, an eleven DOFs simplified model has been developed, which contains the vertical movements of the front and rear wheels, the longitudinal, vertical and pitch movements of car body, and the longitudinal, lateral, vertical, roll, pitch, and yaw movements of the powertrain block. The vehicle coordinate system defined in Figure 7 and the powertrain local coordinate system shown in Figure 3 are consistent. The parameters of stiffness and damping are obtained by test, and all parameters of this model are given in Table 3 of Appendix 2. The vector of generalized coordinates is defined in the following matrix form as q \u00bc xP yP zP P P P xCB zCB CB zFW zRW T \u00f017\u00de where x, y, z, , , and indicate, respectively, the longitudinal, lateral, vertical, roll, pitch and yaw displacements of the rigid block; indications of italic subscripts P, CB, FW, and RW are also listed in Appendix 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001143_acc.2013.6579933-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001143_acc.2013.6579933-Figure1-1.png",
        "caption": "Figure 1. Swashplate Mechanism",
        "texts": [
            " The modeling process is briefly outlined and followed by output and input variance constrained control design. Extensive numerical investigations of closed loop systems eigenvalues variations and time responses indicate that these controllers are robustly stable for modeling uncertainties in flight speeds, inertial properties, initial conditions and noise intensities. I. INTRODUCTION Helicopter control is a difficult problem. Technologically, aerodynamic forces and moments are generated using a sophisticated device called the swashplate (Fig. 1) which controls blade pitch angles. Blades\u2019 motion must be correlated to achieve different pitch angles to compensate the \u201cdissymmetry of lift\u201d [1] and ensure straight level flight by counter-acting the tendency to roll and yaw if the blades have the same angle of attack. Blades\u2019 attachments to the rotor hub must be soft to reduce the stress level and allow two fundamental motions: flapping and lead-lagging (Figs. 2, 3). Flapping (up-down motion) modulates the lift force according to the blade position in the rotor hub plane and alleviates the dissymmetry of lift, while lead-lagging (forward-backward motion) reduces the ground effect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003113_s0146411620020078-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003113_s0146411620020078-Figure9-1.png",
        "caption": "Fig. 9. Test mission 2 3D trajectory.",
        "texts": [
            " To prove the capability of the designed controller integrated with the CC path-following algorithm, one of the performed scenarios is demonstrated herein where a change of height is included to simulate the process of takeoff and landing sequence in addition to a constant height path segment with both straight line and loitering actions. The f light conditions in this scenario include 100% crosswind all the AUTOMATIC CONTROL AND COMPUTER SCIENCES Vol. 54 No. 2 2020 time with the speed of 10 [m/s] and model uncertainties as described earlier. Figure 9 shows the performed trajectory that includes ascending and descending and also two loitering points. The integrated controller and path following algorithm can easily switch between the requested modes in a very robust behavior. For this scenario, at a constant altitude segment (h = 200 [m]), the pitch channel control law regulates the desired trim altitude for the Tiger\u2013Trainer UAV. Meanwhile, the roll channel control law responds to the desired bank angle to achieve the required turning. The altitude hold control loop keeps the altitude constant despite the change of the bank angles to turn in responding to different sharp maneuvers as shown in Fig. 9 which also illustrates the robustness of the proposed f light control model against wind disturbance. In this work, a detailed analysis of Carrot Chasing autonomous path-following guidance algorithms was performed with NDI, INDI and MINDI flight control systems. A small UAV nonlinear 6-DOF rigid body model is used. NDI is a powerful f light controller for the accurate model by designing asymptotic stable control law. However, the accurate model is unavailable, therefore a modification was required to increase the f light control robustness against wind disturbances and model uncertainties"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002233_1350650116661071-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002233_1350650116661071-Figure11-1.png",
        "caption": "Figure 11. Kinematics relation of the cam and the tappet.",
        "texts": [
            " The PDF of the asperity heights of the nominally elastic rough surface of the cam\u2013tappet pair in the studied valve train is \u00f0z\u00de \u00bc 1ffiffiffiffiffi 2 p 0 exp z2 2 20 \u00f021\u00de where 0\u00bc 0.47 mm. As the wear coefficient value of the cam material under boundary lubrication conditions is obtained in the \u2018\u2018Measuring wear coefficient of cam material in boundary lubrication\u2019\u2019 section, which is kb\u00bc 4.81 10 11mm3N 1mm 1, the wear coefficient for mixed lubrication contacts, k, can be calculated using equation (4), and its value varies with the oil film thickness is shown in Figure 10. Calculation of the relative sliding velocity In Figure 11, the cam rotates counter clockwise at a speed of !c and contacts the tappet at point P (e1, e2). The following equations are obtained from the kinematics analysis e1 \u00bc dlP=d e2 \u00bc r0 \u00fe lP \u00f022\u00de at University of Nottingham on July 29, 2016pij.sagepub.comDownloaded from where lp is the lift of the contact point. In the fixed reference frame n1on2, the velocity of the contact point Pc on the cam is given by ~vPc \u00bc e2!c~n1 \u00fe e1!c~n2 \u00f023\u00de The velocity of the contact point Pf on the tappet is given by ~vPf \u00bc e1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002190_tmag.2016.2589924-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002190_tmag.2016.2589924-Figure1-1.png",
        "caption": "Fig. 1. SPMPM machine and flux lines by FEM.",
        "texts": [
            " Finally, ACHFO cannot determine the amplitude and the phase of space orders [10]. Although the tool gives a complete picture of each harmonic of the radial pressure versus time and space orders, some of them have no influence on the magnetic noise [11]\u2013[13]. In order to validate the origin of the radial pressure, different approaches are compared such as the convolution approach [14]. In addition, an experimental operational deflection shape (ODS) measurement is performed to show the deflection shape of the low space order. The structure is presented in Fig. 1, a modular machine defined by 2 p = Zs \u00b1 2, (Zs is number of slots fixed to Zs = 12 and pole pair number p = 5). What is more, this machine is characterized by a winding factor Kw(1) = 0.933, Spp = 0.4, least common multiple (lcm) = 60, and greatest common divider (gcd) = 2. The winding factor Kw and the number of slots per pole per 0018-9464 \u00a9 2016 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_30_0001389_sii.2014.7028020-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001389_sii.2014.7028020-Figure1-1.png",
        "caption": "Fig. 1. Two-Wheeled Robot with IR sensors and an Omni-Vision Camera",
        "texts": [
            " Typical examples are flocking in groups of birds, schooling in groups of fish, and herding in animals. Biologists have studied such self-organized pattern formations for many years. Lukeman et al gathered and analyzed a highquality dataset of flocking surf scoters [10]. A preference for neighbors directly in front was revealed by generating spatial and angular neighbor-distribution plots. Ballerini et al reported that each bird interacts on average with six or seven neighbors, rather than with all neighbors within a fixed metric distance. This interaction is a proven anti-predatory mechanism. Figure 1 shows the two wheeled mobile robot used in this paper. The robot is 30 cm tall and 17 cm wide, and equipped with five distance sensors and an omnidirectional camera. Its input-output cycle time is approximately 0.75 s. Figure 2 shows the range of each behavioral rule, including frontal interaction. The repulsion and attraction rules enable the robot to avoid collisions and remain close to its neighbors, respectively. The direction-matching rule is 978-1-4799-6944-9/14/$31.00 \u00a92014 IEEE 106 a pseudo-alignment rule based on the relative positions of an individual\u2019s neighbors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002721_s10338-019-00156-w-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002721_s10338-019-00156-w-Figure6-1.png",
        "caption": "Fig. 6. Simplified bolted joint beam with virtual material",
        "texts": [
            " 5, the bolted beam system consists of two 300\u00d760\u00d7103 simple sheet beams connected by two O8 bolts, and the material is low-carbon steel with Young\u2019s modulus E = 210e11 Pa, Poisson\u2019s ratio \u03c5 = 0.3 and density \u03c1 = 7.8 \u00d7 103 kg/m3. An approximate free\u2013free boundary condition is simulated by suspending the beam at both ends from the ceiling. The input excitations applied by the model hammer are measured by the integral transducer at point A, while the accelerations are measured by the miniature single-axial accelerometers at point B. The simplified finite element model for the bolted joint beam is shown in Fig. 6. The joint interfaces are replaced by the equivalent virtual material. The material properties are defined as the optimization variable collection: {Young\u2019s modulus E, Poisson\u2019s ratio \u03c5, density \u03c1}. Based on the genetic algorithms for multi-objective optimization, the experimental results of the first three modal frequencies are chosen as the objective functions to identify the mechanical properties of the virtual material, i.e., {E, \u03c5, \u03c1}. To minimize the difference between the calculated modal frequencies and the experimental results, three objective functions and three variable parameters are defined as minimize: \u03d5i(X ) = \u2016\u03c9sim i \u2212\u03c9exp i \u2016 \u2016\u03c9exp i \u2016 , i = 1, 2, 3 subject to: X = [E, \u03c5, \u03c1]T (15) where the superscript \u201csim\u201d refers to simulation results, and \u201cexp\u201d represents experimental results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure3-1.png",
        "caption": "Fig. 3. Conventional 6S-10P E-Core SFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " To overcome PM usage in SFPMM, an overview is carried out in [6] showing different topology of SFPMM with reduced PM usage whereas E-core SFPMM is introduced in [11] which reduced PM usage to half and term as conventional 12/10 stator slot/rotor pole (12S-10P E-Core SFPMM) as shown in Fig. 2(a). However, flux leakage is neglected and not taken in consideration. Flux distribution of 12S-10P E-Core SFPMM is shown in Fig. 2(b). Analysis reveals that despite of reduce in PM usage, still there are significant flux leakage. A similar performance evaluation of conventional 6S-10P SFPMM with E-Core and C-Core is carried out in [12] that reduce PM usage and enhance slot area. Moreover, its performance is compared with conventional 12S-10P E-Core SFPMM. Fig. 3(a) and Fig. 3(b) shows 2-D cross sectional view and flux distribution of 6S-10P E-Core SFPMM whereas Fig. 4(a) and Fig. 4(b) shows 2-D cross sectional view and flux distribution of 6S-10P C-Core SFPMM. Analysis reveals that conventional 6S-10P E-Core SFPMM and 6S-10P C-Core SFPMM retain same PM volume as that of 12S-10P E-Core SFPMM and increased slot area, however author fails to compensate effects of leakage flux. Until now, many researchers tried to reduced PM as much as possible and suppress flux leakage but unfortunately both effects are not considered at a time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure24-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure24-1.png",
        "caption": "Fig. 24 Discretized model using mesh options mesh 1 and mesh 4 as given in Table 1: (a) mesh 1 and (b) mesh 4",
        "texts": [],
        "surrounding_texts": [
            "[1] Thurm, U., Erler, G., Godde, J., Kastrup, H., and Keil, T., 1983, \u201cCilia Special- ized for Mechanoreception,\u201d J. Submicrosc. Cytol., 15(1), pp. 151\u2013155. [2] Keil, T. A., and Steinbrecht, R. A., 1984, \u201cMechanosensitive and Olfactory Sensilla of Insects,\u201d Insect Ultrastructure, Vol. 2, R. C. King and H. Akai, eds., Plenum Press, New York, pp. 477\u2013516. [3] Keil, T. A., 1997, \u201cFunctional Morphology of Insect Mechanoreceptors,\u201d Microsc. Res. Tech., 39(6), pp. 506\u2013531. [4] Palka, J., Levine, R., and Schubiger, M., 1977, \u201cCercus-to-Giant Interneuron System of Crickets. 1. Some Attributes of Sensory Cells,\u201d J. Comp. Physiol., 119(3), pp. 267\u2013283. [5] Shimozawa, T., and Kanou, M., 1984, \u201cVarieties of Filiform Hairs: Fractionation by Sensory Afferents and Cercal Interneurons of a Cricket,\u201d J. Comp. Physiol., 155(4), pp. 485\u2013493. [6] Jacobs, G. A., Miller, J. P., and Aldworth, Z. A., 2008, \u201cComputational Mechanisms of Mechanosensory Processing in the Cricket,\u201d J. Exp. Biol., 211(11), pp. 1819\u20131828. [7] Gnatzy, W., and Tautz, J., 1980, \u201cUltrastructure and Mechanical Properties of an Insect Mechanoreceptor: Stimulus-Transmitting Structures and Sensory Apparatus of the Cercal Filiform Hairs of Gryllus,\u201d Cell Tissue Res., 213(3), pp. 441\u2013463. [8] Landolfa, M. A., and Miller, J. P., 1995, \u201cStimulus-Response Properties of Cricket Cercal Filiform Receptors,\u201d J. Comp. Physiol., A., 177(6), pp. 749\u2013757. [9] Magal, C., Dangles, O., Caparroy, P., and Casas, J., 2006, \u201cHair Canopy of Cricket Sensory System Tuned to Predator Signals,\u201d J. Theor. Biol., 241(3), pp. 459\u2013466. [10] Steinmann, T., Casas, J., Krijnen, G., and Dangles, O., 2006, \u201cAir-Flow Sensitive Hairs: Boundary Layers in Oscillatory Flows Around Arthropod Appendages,\u201d J. Exp. Biol., 209(21), pp. 4398\u20134408. [11] Cummins, B., Gedeon, T., Klapper, I., and Cortez, R., 2007, \u201cInteraction Between Arthropod Filiform Hairs in a Fluid Environment,\u201d J. Theor. Biol., 247(2), pp. 266\u2013280. [12] Dangles, O., Steinmann, T., Pierre, D., Vannier, F., and Casas, J., 2008, \u201cRelative Contributions of Organ Shape and Receptor Arrangement to the Design of Crickets Cercal System,\u201d J. Comp. Physiol., A, 194(7), pp. 653\u2013663. [13] Heys, J., Gedeon, T., Knott, B. C., and Kim, Y., 2008, \u201cModeling Arthropod Hair Motion Using the Penalty Immersed Boundary Method,\u201d J. Biomech., 41(5), pp. 977\u2013984. [14] Casas, J., and Dangles, O., 2010, \u201cPhysical Ecology of Fluid Flow Sensing in Arthropods,\u201d Annu. Rev. Entomol., 55(1), pp. 505\u2013520. [15] Cummins, B., and Gedeon, T., 2012, \u201cAssessing the Mechanical Response of Groups of Arthropod Filiform Flow Sensors,\u201d Frontiers in Sensing: From Biology to Engineering, F. G. Barth, J. A. C. Humphrey, and M. V. Srinivasan, eds., Springer, Wien, New York, pp. 239\u2013250. [16] Czaplewski, D. A., Ilic, B. R., Zalalutdinov, M., Olbricht, W. L., Zehnder, A. T., and Craighead, H. G., 2004, \u201cA Micromechanical Flow Sensor for Microfluidic Applications,\u201d J. Microelectromech. Syst., 13(4), pp. 576\u2013585. [17] Krijnen, G., Lammerink, T., Wiegerink, R., and Casas, J., 2007, \u201cCricket Inspired Flow-Sensor Arrays,\u201d IEEE Sensors Conference, Atlanta, GA, Oct. 28\u201331, pp. 539\u2013546. [18] Casas, J., Liu, C., and Krijnen, G. J. M., 2012, \u201cBiomimetic Flow Sensors,\u201d Encyclopedia of Nanotechnology, B. Bhushan, ed., Springer, The Netherlands, pp. 264\u2013276. [19] Dagamseh, A. M. K., Wiegerink, R. J., Lammerink, T. S. J., and Krijnen, G. J. M., 2012, \u201cTowards a High-Resolution Flow Camera Using Artificial Hair Sensor Arrays for Flow Pattern Observations,\u201d Bioinspiration Biomimetics, 7(4), p. 046009. [20] Droogendijk, H., Casas, J., Steinmann, T., and Krijnen, G. J. M., 2015, \u201cPerformance Assessment of Bio-Inspired Systems: Flow Sensing MEMS Hairs,\u201d Bioinspiration Biomimetics, 10(1), p. 016001. [21] Edwards, J. S., and Palka, J., 1974, \u201cCerci and Abdominal Giant Fibers of House Cricket, Acheta-Domesticus.1. Anatomy and Physiology of Normal Adults,\u201d Proc. R. Soc. London, Ser. A, 185(1078), pp. 83\u2013103. 081006-10 / Vol. 138, AUGUST 2016 Transactions of the ASME Downloaded From: http://biomechanical.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jbendy/935362/ on 03/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use [22] Miller, J. P., Krueger, S., Heys, J. J., and Gedeon, T., 2011, \u201cQuantitative Characterization of the Filiform Mechanosensory Hair Array on the Cricket Cercus,\u201d PLoS One, 6(11), p. e27873. [23] Bitplane, 2014, \u201cImaris,\u201d Bitplane USA, Concord, MA. [24] Joshi, K., 2012, \u201cBiomechanical Analysis of a Cricket Filiform Hair Socket Under Low Velocity Air Currents,\u201d M.S. thesis, Montana State University, Bozeman, MT. [25] Shimozawa, T., Kumagai, T., and Baba, Y., 1998, \u201cStructural Scaling and Functional Design of the Cercal Wind-Receptor Hairs of Cricket,\u201d J. Comp. Physiol., A, 183(2), pp. 171\u2013186. [26] Vincent, J., and Wegst, U. G. K., 2004, \u201cDesign and Mechanical Properties of Insect Cuticle,\u201d Arthropod Struct. Dev., 33(3), pp. 187\u2013199. [27] M\u20aculler, M., Olek, M., Giersig, M., and Schmitz, H., 2008, \u201cMicromechanical Properties of Consecutive Layers in Specialized Insect Cuticle: The Gula of Pachnoda marginata (Coleoptera, Scarabaeidae) and the Infrared Sensilla of Melanophila acuminata (Coleoptera, Buprestidae),\u201d J. Exp. Biol., 211(16), pp. 2576\u20132583. [28] Dechant, H. E., Rammerstorfer, F. G., and Barth, F. G., 2001, \u201cArthropod Touch Reception: Stimulus Transformation and Finite Element Model of Spider Tactile Hairs,\u201d J. Comp. Physiol., A, 187(4), pp. 313\u2013322. Journal of Biomechanical Engineering AUGUST 2016, Vol. 138 / 081006-11 Downloaded From: http://biomechanical.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jbendy/935362/ on 03/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_30_0003789_cac51589.2020.9327203-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003789_cac51589.2020.9327203-Figure1-1.png",
        "caption": "Fig. 1. Quadrotor UAV manipulator system with coordinate reference frames.",
        "texts": [
            " The multiplication between quaternions q and \u03c3, q \u2297 \u03c3 = [ q0\u03c30 \u2212 qT v \u03c3v q0\u03c3v + \u03c30qv \u2212 S(\u03c3v)qv ] (1) Moreover, CB A , representing the relationship between coordinate frames A and B, can be described by q as, CB A = (q20 \u2212 qT v qv)I3 + 2qvq T v + 2q0S(qv) (2) where S(\u00b7) is a operator for a vector, e.g., \u03ba =[ \u03ba1 \u03ba2 \u03ba3 ]T , to a skew symmetric matrix, and it can be described by, S(\u03ba) = \u23a1\u23a3 0 \u2212\u03ba3 \u03ba2 \u03ba3 0 \u2212\u03ba1 \u2212\u03ba2 \u03ba1 0 \u23a4\u23a6 (3) This part briefly provides the model of UAV manipulator attitude system according to our previous work [11]. As depicted in Fig. 1, some coordinate reference frames are given by, OI : the world-fixed inertial reference frame. Ob: the body-fixed reference frame. Oi: the frame of robotic arm i, in which (i = 1, 2) is the link number. Additionally, some symbol definitions are provided by, pI = [x, y, z]T : the position of Ob with reference to OI . \u03a8 = [\u03d5, \u03b8, \u03c8]T : Euler angles for definition of the UAV attitude. \u03b7 = [\u03b71, \u03b72] T : robotic arm joint angles. H = [(pI)T ,\u03a8T ,\u03b7T ]T \u2208 R8\u00d71: the vector with all generalized variables included"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000844_1.4027398-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000844_1.4027398-Figure5-1.png",
        "caption": "Fig. 5 Schematic diagram of journal and bearing",
        "texts": [
            " The rotational speed, torque, and power of the engine are collected and processed by the testing system of the engine testing bench. The measuring method of the 3D orbit of the journal center of the crankshaft bearing is that the locations of the journal center in the cross section of the bearing and at the axial direction of the bearing are measured, respectively, at the same time, and then the data in one operating condition of the engine are processed synthetically to obtain the 3D orbit. The method to obtain the location of the journal center in the cross section of the bearing is as follows. As shown in Fig. 5, the oil film thickness h1 and h2 are measured by two eddy current gap sensors mounted in the main bearing cap, then the x and z coordinates of the journal center location Oj in the bearing are calculated by solving Eq. (1): \u00bd\u00f0Rb h1\u00de sin a1 \u00fe x 2 \u00fe \u00bd\u00f0Rb h1\u00de cos a1 \u00fe z 2 \u00bc R2 j \u00bd\u00f0Rb h2\u00de sin a2 x 2 \u00fe \u00bd\u00f0Rb h2\u00de cos a2 \u00fe z 2 \u00bc R2 j (1) where Rb is the bearing radius, Rj is the journal radius, and a1 and a2 are the angles between the axis of two eddy current sensors and the negative direction of the z coordinate axis, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure3-1.png",
        "caption": "Figure 3. Five circles on the surface of OCT.",
        "texts": [
            " Either the angle j j is small enough to satisfy Sinj j4 b/a or j j is such as Sinj j> b/a; there is no angle j j satisfying a2Sin2 \u00bc b2Sin2 . Consequently, when the given system (a, b, j j) of two lengths and one angle is appropriate, there are four RR open chains producing the same OCT for a point M of the outermost body in the triplet of rigid bodies. The metric properties of the four chains are symbolized by (b, ; a), (a, ; b), (b, ; a), and (a, ; b) where the combination (b, ; a) characterizes the original mechanism used to obtain the OCT equation. In Figure 3, through an arbitrary point Q on the surface of the OCT, five circles drawn to the OCT surface intersect. They are the parallel of latitude and four circular generatrices. One can state a theorem of Euclidean geometry: for a broad family of OCTs, an OCT has four circular generatrices, which lie on planes containing the OCT center. From the datum of one circular generatrix of an OCT, the other three circular generatrices can be derived. That theorem is complemented by the property: any plane of slope, either j j or j j and passing through the OCT center that is the origin O, intersects the surface along a couple of congruent circles",
            " The bar (OB2) is constrained to rotate with respect to (OB1) by a revolute pair whose axis (O, k) is perpendicular to the two bars (OB1) and (OB2). The bar (MB2) can rotate with respect to (MB1) around any axis passing through the point M that is a center of spherical motion. Due to the permanent line-symmetry of the (OB1MB2) isogram, the bar (MB2) rotates with respect to (MB1) around an axis passing through the point M and being perpendicular to the bars (MB1) and (MB2). Consequently, an R pair can be placed at M in replacement of the S pair as described in the self-explanatory Figure 10, which stems from Figure 3. The over-constraining R pair axis passes through M and is perpendicular to the adjacent bars (MB1) and (MB2). The obtained 4R chain is a Bennett linkage. Each link has its opposite, which is congruent to itself because a line-symmetry is a particular orientation-preserving isometry. This is why the conventional orientation of twist angles used to study OCTs has to be modified in order to become consistent with the line symmetry. The vector (B1M) is linesymmetric with (B2O) which is equal to (OB2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001020_robio.2014.7090466-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001020_robio.2014.7090466-Figure1-1.png",
        "caption": "Fig. 1. Transporting a rod orthogonally to the long edge.",
        "texts": [
            " As a result, the object can be linearly transported to a target area, even if the object has an irregular shape. The approach is effective for objects with some volume such as regular polygons, but does not work well for objects with a long and narrow shape. For example, consider an object such as a rod. Such an object has restricted pushable edges, and therefore, once it inclines, it is difficult to be transported while suppressing its rolling. In this case, it would rather easy to be transported such that the long pushable edge has a direction orthogonal to the direction to the goal as shown in Fig. 1. In this paper, we propose another mobile agent based approach for transporting an object while rolling the object whose long edge is orthogonal to the direction to a goal. The structure of the balance of this paper is as follows. In the second section, we describe the background. The third section describes our transport model based on pheromone agents. In the fourth section, we the numerical experiments using a simulator based on our algorithm. Finally, we conclude our discussions in the fifth section and present future research directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000118_s11071-019-05341-7-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000118_s11071-019-05341-7-Figure1-1.png",
        "caption": "Fig. 1 An underactuated system with time-varying weak coupling",
        "texts": [
            " (ii) For any initial state xm(0), we compute \u03b1m(0) = xm(0) \u2212 \u03c6m(0) and choose \u03b2m,0 such that |\u03b1m(0)| < \u03b2m,0 where m = 2, . . . , n \u2212 1. Then, we determine the virtual control law \u03c6m+1 = \u2212\u03b8m\u03b3m with \u03b8m > 0. (iii) Given any initial state xn(0), we find en(0) = xn(0) \u2212 vn(0) and choose \u03b2n,0 such that |\u03b1n(0)| < \u03b2n,0. Finally, we determine the actual control law u = \u2212 \u03b8n\u03b3n with \u03b8n > 0 in (7). The underactuated, weakly coupled, and unstable mechanical system presented in [1] is considered as a practical simulation example. As shown in Fig. 1, a mass\u2013spring system with a mass m1 on a horizontal smooth surface is interconnected to an inverted pendulum m1 connected by a nonlinear spring where the control force only acts on the mass m1. In Fig. 1, z1 \u2208 (\u2212\u03c0/2, \u03c0/2) is the angle of the pendulum from the vertical, z3 is the displacement ofmass of themass\u2013 spring system, kd is the spring coefficient for the nonlin- ear spring, kw denotes the spring coefficient for the linear spring, g is the gravitational acceleration, l denotes the pendulumheight,m1 is themass of themass\u2013spring system, and m2 denotes the mass of the inverted pendulum. For more practical application, it is assumed that a nonlinear spring has the potential performance deterioration of hardening spring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002742_s42235-020-0009-4-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002742_s42235-020-0009-4-Figure2-1.png",
        "caption": "Fig. 2 Magnetic tweezers system CAD design and explosive view.",
        "texts": [
            " Section 4 introduces the magnetic force generation ability of the system using the experimental data shown in section 3 and the conclusions are in section 5. The system setup shown in Fig. 1 illustrates the connection and signal transfer between each of the hardware components comprising the magnetic tweezers. Each of the three GW Instek APS 1102A programmable digital AC/DC power supplies connects to a pair of coils on the magnetic tweezers through a National Instrument SCB-68A connector block and AXICOM D3023 Relay. A CAD design of the hexagonal magnetic yokes can be seen in Fig. 2. These were 3D printed using Proto-pasta Magnetic Iron PLA (0.15 T magnetic saturation) and have outer and inner side to side lengths of 115.3 mm and 71.3 mm, respectively. The magnetic yokes function to create a closed magnetic circuit when the system is activated, which reduces the excitation current necessary for generating a magnetic field[4] and prevents excess heat generation within the system. The visual acquisition was developed using a Pixelink D734CU-T color camera configured to work under 6 Frame Per Second (FPS) with a resolution of 992 \u00d7 992 pixels",
            " All the six electromagnetic coils were fabricated using AWG-25 heavy-built insulation coated copper wire with 527 turns per each coil, which can each produce 790 Ampere-turns under a maximum input current of 1.5 Amps for magnetic field gradient generation. Our group used cobalt iron alloy (VACOFLUX 50, VACUUMSCHMELZE GmbH & Co.KG) material to fabricate the poles, which were laser-cut to shape by Polaris Laser Laminations, LLC. The average dimension of the sharp tip on each pole were within a radius of 40 \u03bcm[6]. All six poles were installed on the yokes, with three poles per yoke layer to form the actuation coordinate system XaYaZa which can be seen in Fig. 2. Through coordinate system transformation, we can transform the arbitrary coordinates we set in the measurement coordinate system XmYmZm to XaYaZa for our control strategy. All the location information in experimental results and analysis in this paper are consistent with the measurement coordinate system in Fig. 2. Vertical distance between poles on top and bottom planes is 2.04 mm. Poles on the same plane have a gap of 3.67 mm between each other and the effective working space in the center of the system is about 2.5 mm \u00d7 2.5 mm \u00d7 0.5 mm[6]. The closed-loop control interface was created using a MATLAB graphical user interface. Fig. 3 illustrates the flow of the control algorithm. Images captured by the camera were processed to get the location and area in- formation of the target microswimmers (Fabrication process is shown in section 3)",
            " (10) with data obtained from the experiments. From section 3.2, it\u2019s known that the 3D control of trajectory \u2018\u2018M\u2019\u2019 is the most stable one, thus our calculation and analysis proceeded using this specific sample. The black dashed vertical lines in each graph of Fig. 7 represents a change in the desired location. Fig. 7a shows the current history during the whole experiment period, in which the current on the pair of coils in xa was led in the negative direction for most of the time. By looking at the coordinate system in Fig. 2 and comparing it with Fig. 6 for \u2018\u2018M\u2019\u2019, it is easy to find out that the main moving direction of the microswimmer has a direction component along negative xm, after coordinate transformation, the force vector in xa is also negative, and the other two terms will change sign depending on the ongoing moving directions. The current, velocity and force magnitudes drop at the dashed lines. This is because the microswimmer approached the desired location and maintained position with an exciting current power in Journal of Bionic Engineering (2020) Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000991_s00170-015-6812-0-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000991_s00170-015-6812-0-Figure15-1.png",
        "caption": "Fig. 15 Energy-dispersive X-ray spectrum along a the depth direction and b the width direction of the unit",
        "texts": [
            "6 Microstructure The microstructure of the units after laser remelting by different parameters is alike. Figure 14 shows the unit\u2019s microstructure obtained with electric current 150 A, pulse duration 10 ms, frequency 5 Hz, defocus distance 6.5 mm, and scanning speed 0.5 mm/s. Figure 14a, b shows the microstructure of the cross section of the transition zone after laser remelting. The transition zone is composed of a dendritic structure and the substrate microstructure is tempered martensite and granular carbide. Figure 15c shows the microstructure of the melted area. It can be seen that the microstructure of the melted zone is a very fine dendritic structure. Energy-dispersive X-ray (EDX) spectrum was used to obtain the element composition. As shown in Fig. 14d, the interdendritic space is filled by a carbide network (C, Cr, V, Si, Mo) and the weight and atomic percentages are listed. In the fine dendritic structure, it is filled by elements of Fe, C, Si, V, and Cr as shown in Fig. 14e. Figure 15 is the line scanning image by energy-dispersive X-ray spectrum along the depth and width direction of the unit. It can be seen that the element Fe, Si, V, Mo, and Cr distributions are even along the two directions. 4.7 Microhardness The microhardness of the laser remelting area is shown in Fig. 16. The microhardness changes slightly for different laser parameters. The microhardness achieved in the laser remelting area is higher than that of the H13 matrix. The average microhardness of the melted zone and the transition zone is 620 and 580 HV, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000590_ccdc.2015.7162781-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000590_ccdc.2015.7162781-Figure2-1.png",
        "caption": "Figure 2.the Ground Coordinate System and the Body Coordinate System",
        "texts": [
            " All these methods are based on the accuracy of the modeling. 2 Dynamic Modeling Quad-rotor aircrafts mainly fly low at the near surface, and the range it can reach is limited. So for simplicity, some hypotheses are essential. 1) Ignore the curvature of the earth, and take the earth as a horizontal plane; 2) Take the ground coordinate system as the reference system. Ignore the rotation of the earth and curvilinear motion of the center of mass [10]. The ground coordinate system and the body coordinate system are built, as shown in Figure2. The attitude angles are represented by the Bryant angles. Among them, the pitch angle of the aircraft is defined as\u03b8 , which is the angel between the OX of the body coordinate system and the ground platform and has a range of [-\u03c0 / 2,\u03c0 / 2]. Similarly, the roll angle is the angle between the OZ of the body coordinate system and the vertical plane of the OX, and its range is [-\u03c0 / 2,\u03c0 / 2], too. At last, yaw angle \u03c8 is the angle between the OX projection in the horizontal plane of the body coordinate system and the OX of the ground coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000361_1.4036023-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000361_1.4036023-Figure2-1.png",
        "caption": "Fig. 2 Vector field and limit cycle of Eq. (8) for xn 5 1, f 5 0.5, c 5 1, and d 5 21. The squares indicate the initial states. The diamonds indicate the final state. Both trajectories converge to the same limit cycle, one trajectory starts inside the limit cycle and another starts outside the limit cycle.",
        "texts": [
            " By using the Poincar e\u2013 Bendixson criterion study, the behavior of the oscillator, and the continuously differentiable function V(x)\u00bcx2x2/2 \u00fe _x2/2 as the Lyapunov function, we get f x\u00f0 \u00de rV x\u00f0 \u00de \u00bc x2 x2 n x _x 2fxn _x2 \u00fe c _x2 \u00fe dxx _xffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi _x2 \u00fe x2x2 p (10) The vector field points inward for f(x) rV(x)< 0, outward for f(x) rV(x)> 0, and tangent for f(x) rV(x)\u00bc 0. Evaluating Eq. (10) results in the closed orbit defined by the below equation x2x2 \u00fe _x2 \u00bc c _x \u00fe dxx 2fxn _x \u00fe x2 n x2 x !2 (11) Therefore, there exists a closed orbit in the system; consequently, the proposed oscillator (8) is stable. Graphically, the field vector of the oscillator and the existence of the limit cycle can be seen in one example in Fig. 2 for the system xn\u00bc 1 rad/s, f\u00bc 0.5, c\u00bc 1, and d\u00bc 1, where two trajectories are shown: the initial conditions start inside and outside the limit cycle at x\u00bc 1.5, _x\u00bc 0, and x\u00bc 0.1, _x\u00bc 0. Both trajectories converge to the limit cycle. The directions of the vectors inside the limit cycle point outward; and the vectors outside the limit cycle point inward. This nonlinear forcing function can be used to oscillate a physical system. The amplitude of the oscillation can be adjusted by the value of c, and the frequency of the oscillation can be adjusted by the value of d"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000558_icra.2015.7139946-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000558_icra.2015.7139946-Figure1-1.png",
        "caption": "Fig. 1. Testbed Layout with CREST, Imagers, Catheter/Fixture, and EM Ascension TrakStar",
        "texts": [
            " The focus is to use current sensing technologies, including an EPS system, a stereo camera imaging system in place of fluoroscopy, and a predictive catheter configuration model with measured inputs, in an intelligent way. A testbed was constructed to investigate feedback control of continuum robotic manipulators or, more specifically, steerable catheters. The primary objective was to provide a system with a number of sensory sources representative of those currently available to clinicians. This allows the evaluation of localization and control strategies in the context of a realistic clinical environment. The testbed and associated equipment, shown in Figure 1, consists of a catheter manipulation system, measurement systems, and a data acquisition and management system. A durable continuum manipulator prototype was constructed that provides similar actuation and response to commercial catheters while also being relatively easy and inexpensive to build. This manipulator, shown in Figure 2, is constructed from a flexible polymer with four, diametrically opposed, internal lumens. These supporting lumens provide passageways for the control wires, or tendons, made of monofilament wires attached to the distal cap"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure3-1.png",
        "caption": "Fig. 3 The most usable simple planetary gear train, its modified symbol of Wolf and the lever analogy",
        "texts": [
            " Ideal external torques are always in a certain ratio, expressed with the torque ratio t TD min : TD max : TR \u00bc T1 : \u00fet T1 : 1\u00fe t\u00f0 \u00deT1 \u00bc 1: \u00fet : 1\u00fe t\u00f0 \u00de \u00f08\u00de in which always TD min\\TD max\\ TRj j \u00f09\u00de however: \u2022 How many degrees of freedom F (F = 1 or F = 2) the gear train is running with; \u2022 Which shaft is fixed at F = 1 degree of freedom; \u2022 What is the direction of power flow, respectively does the gear train works as a reducer or a multiplier at F = 1 degree of freedom, respectively such as a collecting or as a separating gear at F = 2 degrees of freedom, i.e. as a differential; \u2022 Does the gear works as a single or as a component train along with others in a compound planetary gear train. The application of the alternative method is best illustrated with the two-carrier compound planetary gear trains consisting of two of the most commonly used planetary gear trains, shown in Fig. 3. Using the modified symbol of Wolf, the linking up of the two gear trains can be done in different ways, as shown in Fig. 4 with structural schemes\u2014by two or one compound shafts, where we get threeshafts, respectively four-shafts compound planetary gear train. The most commonly used version is the first with two compound shafts and three external shafts (Fig. 4a). As seen from the figure, one of the compound shafts has no outset\u2014it is an internal compound shaft, so that at both ends two equal sized but opposite direction torques are acting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000927_s0263574714002434-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000927_s0263574714002434-Figure1-1.png",
        "caption": "Fig. 1. The benchmark cooperative robotic system.",
        "texts": [
            "111.164.128 value of the output of the failing actuator u\u0304j (t), so that the system will remain stable and the output y(t) asymptotically tracks the reference signal ym(t) even when actuator failures occur. In order to develop an adaptive control scheme for a general cooperative robotic system, we first consider a two-dimensional cooperative manipulator system as a benchmark system in our study. The system contains two manipulators, which are attached to each side of a rigid platform as shown in Fig. 1. We will design a controller that can guarantee asymptotic tracking of both height h(t) and angle \u03b8(t) of the system. The adaptive control scheme for this sample robotic system can be extended to general cooperative robotic system such a Hexapod system or a humanoid robot, which will be discussed later on. In many systems such as a humanoid robot, adding another leg to the robot for redundancy is not possible without the loss of functionality of the system. In our robotic system model, the redundancy of the system will came from an additional joint in a robotic manipulator. As in Fig. 1, the system uses three actuators q1, q2, q3 to support a rigid platform that links actuator q1 and q3 together. The actuator q2 is added to increase redundancy in the system to compensate for possible actuator failure that could occur on the left side of the platform. In this study we consider three cases of actuator failure patterns: \u2022 Case 1: no actuator failure occurs, that is, \u03c3 (t) = diag{0, 0, 0}, or \u2022 Case 2: the actuator q1 fails, \u03c3 (t) = diag{1, 0, 0}, or \u2022 Case 3: the actuator q2 fails, \u03c3 (t) = diag{0, 1, 0}",
            " Although in this way the outputs of the system will be directly related to each coordinate, the dynamic equations are more complicated because the force of each actuator will act on different directions to each coordinate. On the other hand, by selecting the position of each actuator q1, q2, q3 as the generalized coordinates, we can reduce the complexity of the derivation; as a result, we will use this second set as the generalized coordinates. Before we derive the dynamic equations of the system, we consider the relationship between the coordinates and the outputs of the system. From Fig. 1, we can write the height and the angle of the platform in terms of the position of each actuator as h(q1, q2, q3) = q1 + q2 + q3 2 (6) \u03b8(q1, q2, q3) = arctan ( q1 + q2 \u2212 q3 b0 ) , (7) where the constant b0 is the lenght of the base of the platform. The derivatives of (6) and (7) are h\u0307(q1, q2, q3) = 1 2 (q\u03071 + q\u03072 + q\u03073) (8) \u03b8\u0307(q1, q2, q3) = b0(q\u03071 + q\u03072 \u2212 q\u03073) b2 0 + (q1 + q2 \u2212 q3)2 . (9) Based on the generalize coordinate qi for i = 1, 2, 3, we consider Lagrange\u2019s equation d dt ( \u2202L \u2202q\u0307i ) \u2212 \u2202L \u2202qi = \u03c4i, (10) where the Lagrangian L = T \u2212 V , T is the kinetic energy, V is the potential energy, and \u03c4i is the torque of each actuator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003623_ecce44975.2020.9235618-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003623_ecce44975.2020.9235618-Figure5-1.png",
        "caption": "Fig. 5. 2D FEM model of 18-12 pole SRM.",
        "texts": [
            " The radial force sum of square current and proposed current corresponding to Fig. 3 are shown in Fig. 4. The radial force sum generated by square current has high variation of 616 Np-p. On the other hand, the radial force sum of the proposed current has reduced variation of 158 Np-p. Compared to the radial force sum in Fig. 2 (b), the radial force sum of the proposed current has more variation because the current is optimized to realize not only noise reduction but also reduce the RMS current for efficiency improvement [20] - [21]. Fig. 5 shows the FEM model of the 18-12 pole SRM for electromagnetic analysis. This model uses 2D FEM ANSYS Maxwell for time efficient calculation. The motor performance and the radial force density calculation are carried out for further analysis in the next section. The radial force density is defined by ( , ), where t and \u03b1 are time and spatial angular position. The radial force density ( , ) is calculated by using Maxwell Tensor expressed in [14] [16] as, ( , ) = ( \u2212 ), (4) where Br is the radial magnetic flux density, Bt is the tangential magnetic flux density, and is the permeability of free space",
            " Therefore, it can be concluded that the 18-12 pole SRM has deformation with significant mode 0. Fig. 9 (a) and (b) show the radial force density excited by square current and proposed current in one period of electrical cycle, respectively. These force densities are carried out at operating point of 1000 rpm and 10 Nm. The radial force density is obtained by using the previously shown equation of (4). The 0\u00b0 of spatial angle is defined at the positive x-axis and y-axis of coordinate as previously shown in Fig. 5. Both of square and proposed current have six spatial force density waveforms, which is correspondent to the number of stator teeth for each phase. The color scale provides us an idea as follows; (i) the radial force density patterns are similar in Fig. 9 (a) and (b), (ii) the density is high in Fig. 9 (a). To analyze the effectiveness of radial force sum flattening in the 18-12 pole SRM, the FFT is applied to the radial force densities of the square and proposed currents with respect to the spatial angular position in a constant time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000382_ijhvs.2018.089897-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000382_ijhvs.2018.089897-Figure2-1.png",
        "caption": "Figure 2 Three-dimensional trailer model",
        "texts": [
            " Section 2 introduces the mechanism of the three-dimensional trailer model with the characteristics mentioned earlier; Section 3 presents the static analysis of the proposed model; the results presented in Section 3 are analysed and discussed within a case study in Section 4; and finally, conclusions are drawn in Section 5. Taking into consideration that the last trailer of LCVs is the critical unit, the model of trailer is simplified (Figure 1), and a mechanical system that represents this unit is developed (Figure 2). Mechanical systems can be represented by kinematic chains composed of links and joints, which facilitates their modelling and analysis (Kutzbach, 1929; Crossley, 1964; Tsai, 2001). Using mechanism theory, a three-dimensional model that represents the last trailer is proposed (Figure 2). More details concerning the developed model are contained in the technical report to this paper, which includes the modelling and analysis of the trailer model (Moreno et al., 2016c). The model is composed of three mechanisms: \u2022 the first mechanism is located at the front of the trailer and is composed of sub-mechanisms that represent the tyres (tyres system), the suspension (rigid suspension system) and the fifth-wheel (fifth-wheel system \u2013 FW) \u2022 the second mechanism is located at the rear of the trailer, and is composed of sub-mechanisms that represent the tyres (tyres system) and the suspension (rigid suspension system) \u2022 the third mechanism represents the trailer body (chassis), and links the front and rear trailer mechanisms"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002650_usys.2016.7893918-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002650_usys.2016.7893918-Figure1-1.png",
        "caption": "Fig. 1. A kinematic diagram of nonholonomic autonomous surface vehicle agent",
        "texts": [
            " Despite the restricted agent connectivity, formation control of the multiple ASV agents is successfully achieved. This paper is organised as followed: Section 2 discusses the preliminaries and problem formulation of the ASV are presented. Section 3 presents the proposed cooperative formation control algorithm. Section 4 and Section 5 presents simulation result to prove its principle workability. We consider first the simple kinematic modeling for a single generic nonholonomic ASV agent (shown in Figure 1) which can be described by \ud835\udc65\ud835\udc56 = \ud835\udc63\ud835\udc56\ud835\udc50\ud835\udc5c\ud835\udc60(\ud835\udf03\ud835\udc56) \ud835\udc66\ud835\udc56 = \ud835\udc63\ud835\udc56\ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udf03\ud835\udc56) \ud835\udf03\ud835\udc56 = \ud835\udf14\ud835\udc56 (1) where \ud835\udc65\ud835\udc56 and \ud835\udc66\ud835\udc56 represent the position of a single ASV agent \ud835\udc56 in an inertial coordinate system, \ud835\udf03\ud835\udc56 represent the orientation of the ASV agent \ud835\udc56, \ud835\udc63\ud835\udc56 represent the forward velocity of the single agent robot \ud835\udc56, and \ud835\udf14\ud835\udc56 represent the angular velocity of agent \ud835\udc56. In this system, the velocity, \ud835\udc63\ud835\udc56 and angular velocity, \ud835\udf14\ud835\udc56 are considered as input control signal to be designed. B. Inter-agent Communication Modeled by a Graph The information about the state of the ASVs agent \ud835\udc56 is available to itself as well as to its neighboring agents subjected to the inter-agent communication connectivity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002508_s11432-015-0153-2-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002508_s11432-015-0153-2-Figure2-1.png",
        "caption": "Figure 2 Working principle of a small-scale helicopter.",
        "texts": [
            " The wireless communication modem is used to monitor the status of the helicopter. In this work, a hybrid model is used to describe the dynamics of a small-scale helicopter, which includes the nonlinear rigid body dynamics, main rotor dynamics, and simplified yaw dynamics. The model uncertainties and other trivial factors are treated as external disturbances. The position of the gravity center of the helicopter is notated by P = [x, y, z]T in the inertial frame, where the linear velocity of the gravity center in the inertial frame is given by v = [vx, vy, vz ] T (see Figure 2). Control forces and moments originate mainly from the main and tail rotors, controlled by four inputs: lateral and longitudinal cyclic rotor controls \u03b4lat, \u03b4lon; collective pitch input \u03b4col; and tail rotor collective input \u03b4ped. \u03b4lat and \u03b4lon are used to control roll and pitch motions through swashplate tilting. The rotation speed of the main rotor is controlled by an engine governor and it is not considered. The magnitude of the thrust is controlled by changing the collective pitch angle of the main blades through \u03b4col"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002473_j.measurement.2016.11.034-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002473_j.measurement.2016.11.034-Figure1-1.png",
        "caption": "Fig. 1. The experimental setup as developed by Cuny et al. [16] i",
        "texts": [
            " As a validation of the experimental method, a characterisation of the wear occurring during the high-speed contact between an aluminium-based abradable and a Ti-6-Al-4V cutting tool have been realised. Post-mortem analysis were carried out in order to highlight the capacity of the new device to correlate all the components of the interaction force to the effective wear of the abradable material. 2. Experimental setup A ballistic bench is used to perform the high-speed cutting interaction between the abradable specimen and a rigid tool which mimics the shape of an engine fan blade tip (Fig. 1a). The sample is mounted onto a projectile propelled by the sudden expansion of a compressed nitrogen gas and carefully guided toward the cutting tool by a fixed rail inside the bench\u2019s launching tube. In order to avoid a forward gas leakage between the projectile and the launching tube a plastic plug is glued on the projectile back. The projectile velocity, and thus the interaction velocity, is set by the gas pressure. The penetration depth is set by the tool position and the projectile height (Fig. 1b). The projectile is finally stopped without damage into a receiving tank. The cutting tool is clamped by a screw on a piezoelectric force sensor in order to measure the cutting force component Fc of the interaction force. 2.1. Study of the existing uniaxial dynamometer The dynamic behavior of the existing device was first investigated by FEM modeling using the commercial software Abaqus n (a) and the tool-sample measurement configuration in (b). to calculate its rigidity and its eigenfrequencies",
            " The disagreement on the resonance frequency determination between these two approaches was above 6%. In fact, as shown by the first eigenmodes for the tangential and normal directions, the tool fastening was found to be too weak (as shown by Fig. 2c and d). This was essentially due to the cantilever mounting design and to the lack of rigidity of the preloading screw (diameter 8 mm). In addition, this model also shown a lack of stiffness of the device mounting in the tangential and normal directions (Fig. 1e and f). As shown by previous works [12,17] in high-speed contact measurements between an aluminium-based abradable material specimen and a 42CrMo4 steel tool, the cutting force could reach about 1800 N. The coating wear variation as a function of the specimen position in the cutting direction was measured to be about 69 lm. From the numerical model, we obtained that with a force (c) (f) timized parts are shown in (d), (e) and (f). The new experimental device assembly is of the same intensity the displacement of the cutting edge could be about 100 lm in the cutting direction and 70 lm in the normal direction",
            " The employed correction method reshapes the measured force signal and substantially reduce the coupling between the measurement axis. The reference forces and the corrected forces are then found to be in a pretty good agreement. These results highlighted the real benefits of this correction method to extend the bandwidth of a dynamic force device. Some high speed blade-abradable orthogonal cutting tests were carried out to investigate the signal correction influence on the measurement of the cutting forces. The tool edge geometry (i.e. the edge radius r and the thickness e as shown in Fig. 1b), the incursion depth ti and the incursion velocity Vi are the experimental input parameters. The three components of the interaction force in the cutting, tangential and normal directions (Fc; Ft and Fn respectively) are the measurement results. The cutting tool was made of titanium alloy (Ti-6Al-4V) while the tested abradable sample was an aluminium-based abradable material. The gas gun tank where the experimental device takes place was kept in vacuum at a pressure of 5 mbar. Two tests were carried out in the same conditions at a velocity close to 110 m/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001284_j.ijrmms.2013.04.002-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001284_j.ijrmms.2013.04.002-Figure1-1.png",
        "caption": "Fig. 1. Forces acting upon a flyrock fragment during the flight.",
        "texts": [
            " There is a large amount of previous research which suggests various approaches to define safe distance directly from blasting parameters [1\u20133] by calculating flyrock risk [4\u20137], by implementing novel computing techniques such as artificial neural networks [8,9] or by implementing ballistic approach with or without taking drag into account [10,11,12]. However, it is our opinion that ballistic approach incorporating drag is the most precise. Ballistic approach relies on formulation and solution of differential ballistic equation of flyrock fragment flight. The procedure of formulation and solution to these equations is explained in detail in literature [10]. Starting from the condition of equilibrium of forces (Gravity G, Drag D and Lift L) acting upon the flyrock fragment during the flight (Fig. 1) F !\u00bc G !\u00fe D !\u00fe L ! ; or \u00f01\u00de m a!\u00bc G !\u00fe D !\u00fe L ! ; \u00f02\u00de and neglecting the lift force (L) due to small value, differential equations which describe the fragment flight can be written as 09/$ - see front matter & 2013 Elsevier Ltd. All rights reserved. x.doi.org/10.1016/j.ijrmms.2013.04.002 esponding author. Tel.: +381 60 4 807 819; fax: +381 30 421 078. ail address: sstojadinovic@tf.bor.ac.rs (S. Stojadinovi\u0107). m\u20acx\u00bc \u2212Dx \u00bc \u2212 1 2 \u03c1ACDv2 cos\u03b8 \u00f03\u00de m\u20acy\u00bc\u2212mg\u2212Dy \u00bc \u2212mg\u2212 1 2 \u03c1ACDv2 sin\u03b8 \u00f04\u00de where m is the mass of the flyrock fragment (kg); g is the gravitational acceleration (m/s2); \u03c1 is the density of air (kg/m3); A is the cross sectional area of the flyrock fragment (along a plane perpendicular to the direction of flight) (m2); CD is the drag coefficient; v is the flyrock fragment velocity, in the moment of launch v0\u2014flyrock launch velocity (m/s); and \u03b8 is the velocity vector angle; in the moment of launch, \u03b80 is the flyrock launch angle (degree)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure4-1.png",
        "caption": "Figure 4. Velocity contours for standard sizes.",
        "texts": [],
        "surrounding_texts": [
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nFigures 2, 3 show a finite-volume partition of the internal air domain of a field machine.\nTo obtain the distribution fields of flow rates in the body of the field machine, a series of calculations was carried out at different geometric parameters and drum rotation speeds.\nFigures 4 - 8 show the contours of the velocity fields and the velocity vectors for the initial geometric parameters (Figure 1). The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 36.2 m/s. On the left side of the drum, an",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nincrease in the flow rate is observed, due to the narrowing of the space between the drum and the housing wall.\nA slight discrepancy in the display of velocities in the graphs of displaying fields and vectors is due\nto the peculiarities of displaying and visualizing numerical results.",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0002047_s12206-015-1149-z-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002047_s12206-015-1149-z-Figure2-1.png",
        "caption": "Fig. 2. Configuration of bump-type test GFBs for high-speed PM motor: (a) journal GFB; (b) thrust GFB.",
        "texts": [
            "0 mm, respectively. The motor\u2013compressor system is supported by four GFBs: two journal GFBs supporting radial rotor loads and one pair of thrust GFBs supporting axial loads. We note that Ref. [12] detailed the two-stage turbo compressor section. The PM motor was designed and developed with a rated power of 223 kW (voltage: 300 V; phase current: 570 A) at a maximum rotor speed of 60 krpm for the turbo compressor. Table 1 lists the specification of the analyzed PM motor and its detailed geometry and materials. Fig. 2 shows the typical configuration of the bump-type test GFBs for the use in the high-speed PM motor: (a) journal GFB and (b) thrust GFB. The test GFBs are identical to those in Ref. [12]. The journal GFBs have an axial length of 60 mm and a radius of 35.75 mm. The thrust GFBs have six top-foil pads with an arc angle of 60\u00b0 and inner and outer radii of 27 and 65 mm, respectively. The inclined plane angle at the leading edge, which generates the hydrodynamic film pressure in the air film, is 19.5\u00b0. Table 2 lists the geometry and materials of the bump-type test GFBs, i",
            " Therefore, the temperature effect on the bearing loss is presently neglected. More details on the thermal behavior of the rotor-GFB systems at high temperature are presented in Refs. [10, 16]. On the basis of the validated thermal model of the highspeed PM motor (reference model), two case studies are conducted to investigate the thermal effects of carrier harmonics included in the motor input current and the cooling air supplied by the axial fan at the rotor end. In the first case study, a sine-wave input current where the carrier harmonics are removed (Fig. 2(a) in Ref. [11]) is applied to the reference thermal model with cooling flow. For the second case study, no cooling air from the fan blades in Fig. 3 is assumed to flow to the reference model with the measured input current. Iterative solutions are also obtained between the temperature and electromagnetic losses. Table 5 lists the summary of the calibrated electromagnetic losses, temperatures of the corresponding part, and motor temperatures of the high-speed PM motor at a rotor speed of 35 krpm for the case study with a sine-wave input current and that with no cooling-air supply compared with the reference model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000295_cyber46603.2019.9066742-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000295_cyber46603.2019.9066742-Figure6-1.png",
        "caption": "Figure 6. (a) PLM Welding station and (b) Optimized welding path",
        "texts": [
            "1 , and the number of iterations is 500. Under the specific conditions of the collection, the three algorithms are compared through the path optimization of the floor skeleton solder joints of the car. Table 1 shows the results of the algorithm properties of the basic BAS, basic GA, and BAS-GA in the path optimization process. Figure 2-5 shows the average convergence process and Three-dimensional path comparison, respectively. In this paper, the robot welding station is built using PLM offline programming software, as shown in Figure 6. The planned solder joint path information is imported into the welding workstation, and the above results are simulated in a virtual environment. In the same simulation environment, record the simulation time of the robot to complete the spot welding task. The time required before optimization is 1min41s, and after optimization is 1min15s. The results show that the BAS-GA hybrid optimization algorithm improves the working efficiency of the welding path, which has important guiding significance for the automobile production industry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001084_robio.2014.7090656-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001084_robio.2014.7090656-Figure3-1.png",
        "caption": "Fig 3 the schematic plot of the Experimental set-up",
        "texts": [
            " Furthermore, visual sensing tends to become more accurate as the robot approached the goal. Over the past decade, visual servoing techniques have been applied in the field of mobile robots and many feedback control laws and modeling methods have been proposed [1]. In paper [2] and [3], the visual servoing stabilization of nonholonomic mobile robot with unknown camera parameters is investigated. In these experiences, a pinhole camera fixed to the ceiling is used to measure the robot\u2019s position (Fig 3), but the robot\u2019s orientation is measured by other approaches. In some papers, color recognition [4] is used to obtain a binary image. If in the environments there are other objects with the same color to the target, shape recognition based on Hu moments [9] is also performed. This method can be used to detect the robot, but there is a potential drawback of this method that if in the environments there are other objects not This paper was partially supported by The Scientific Innovation program(13ZZ115), National Natural Science Foundation (61374040), Hujiang Foundation of China (C14002), Graduate Innovation program of Shanghai(54-13-302-102)",
            " The two standard wheels are both driven by Maxon brushless DC motors. The distance between the two motorized standard wheels is 0.5m, and the diameter of the two standard wheels is 0.12m. A camera which is connected to an Intel Pentium IV PC is used to capture image at the rate of 30 fps. Both the camera\u2019s intrinsic and extrinsic parameters are not known. The computer will process the captured image and then to extract specific features. The device\u2019s sampling period is about 13ms per image. The whole experimental set-up is shown in Fig 2, and Fig 3 is its schematic plot. The experimental parameters adopted here are listed in Table 1. We can refer to paper [14] to know these parameters\u2019 explicit definitions. And here has no need to restate them In the experiments, we put a black rectangle board on the robot as a mark (Fig 1), and then on the image plane, take the board\u2019s orientation as the robot\u2019s orientation. First, on the image plane, we use the CAMShift to track the robot, so that we can get a search window that includes the robot only, as is showed in Fig 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.1-1.png",
        "caption": "Figure 3.1 Force on current carrying conductor in magnetic field",
        "texts": [
            "com/go/moritz 36 Electromechanical Motion Systems: Design and Simulation This section reviews the four most popular motors used in electromechanical control systems, detailing modes of operation, limitations and application information for: \u2022 Brush motors \u2022 Brushless motors \u2022 Stepper motors \u2022 Induction motors. The work of Jean Baptiste Biot (1774\u20131862), and Felix Savart (1791\u20131841), formulated the relationship between current in a conductor and the magnetic field surrounding it, which led to Hedrick Antoon Lorentz (1853\u20131928) deriving the formula for the force on a current-carrying conductor in a magnetic field, as illustrated in Figure 3.1. F = BlI (3.1) where B is the magnetic field strength, l is the length of the current-carrying conductor, I is the current magnitude, and F is the Lorentz force. If a flat coil of wire with radius r is placed in a magnetic field and mounted on an axis, as shown in Figure 3.2, then current flowing through the coil, with the polarities shown, will result in Lorentz forces developing on the coil wires, leading to creation of torque and rotation of the coil, where: T = BlIr sin \u03b8 (3.2) As the coil rotates counter clockwise (CCW), the torque will vary in a sinusoidal fashion from a maximum (\u03b8 = 90\u25e6) to zero (\u03b8 = 0\u25e6) at which point motion would stop"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003610_ecce44975.2020.9235657-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003610_ecce44975.2020.9235657-Figure4-1.png",
        "caption": "Fig. 4. Designed in-wheel motor sysytem integrated MMSG and high speed motors.",
        "texts": [
            " Downloaded on December 18,2020 at 23:48:47 UTC from IEEE Xplore. Restrictions apply. can be much decreased by reducing the harmonic fluxes due to pole piece less construction. (iii) It achieves high torque density because the fundamental flux component of the high speed rotor contributes to produce the output torque and it can be obtained by increasing number of the high speed rotor. III. PERFORMANCE EVALUCATION The MMSG and multiple high speed motors were designed to satisfy a demand of 40kW output power for inwheel motor system [12]. Fig. 4 shows the designed MMSG and multiple high speed rotor. As shown in Fig. 4, the gear ratio is 1:20. The maximum output torque of low speed rotor is 500 Nm and the maximum speed of the high speed rotor is 50000 min-1. The MMSG is operated by using the 30 pieces Surface Permanent Magnet Synchronous motor (SPMSM) which output is 1.33kW. The system including the MMSG and these high speed motors is 13.7 kg and the system volume is 4.17\u00d710-3 m3. It realizes the compact and lightweight drive system, and the torque density 120 kNm/m3 and 36.5Nm/kg are achieved. The performances of the mechanical strength, torque density, loss and efficiency in the designed MMSG are evaluated by the FEA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000288_humanoids43949.2019.9035051-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000288_humanoids43949.2019.9035051-Figure2-1.png",
        "caption": "Fig. 2. The experimental setup for vision based joint angles data collection. The camera was positioned at a distance to capture the ArUco markers for estimating the PIP and MCP joint angles of the robot finger. This experiment was conducted for 10 different dynamometer (force-sensor) angles ranging from 0\u25e6 to 90\u25e6 in 10\u25e6 steps. The angle change was achieved by a modular setup that allows a fast adjustment of the dynamometer angle. The experiments conducted involved 10 different trials for every experimental condition (dynamometer angle).",
        "texts": [
            " It must be noted that in certain cases it was observed that the finger continued to exert forces after the reconfiguration stopping point, establishing contact with the fingernail. Such trials were neglected due to the different materials (fingernail is made out of tough resin) and types of physical interaction involved and they were not included in the examined dataset. B. Vision-based Joint Angles Data Collection A second optical setup was created using a 4k web camera and a set of ArUco markers (see Fig. 2). The markers were attached on the finger phalanges instead of the IMU sensors. ArUco is a computer vision processing library developed by Rafael Mu\u00f1oz and Sergio Garrido [21] and it allows the detection of appropriately designed square fiducial markers, providing relative positional data such as the angles and the Cartesian coordinates for each marker. Two markers were attached onto the finger structure (on the two different phalanges) and the dynamometer was mounted on a pivot locking mechanism with angular slots the orientation of which could range from 0\u25e6 to 90\u25e6 in 10\u25e6 steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001450_s11249-013-0247-2-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001450_s11249-013-0247-2-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the experimental setup",
        "texts": [
            ", Yanai, Japan). Before the experiments, the specimens were washed with deionized water and liquid soap and then dried in blowing nitrogen. The temperature and relative humidity in the laboratory were 24 C and 55 %, respectively. 2.2 Equipment Surface roughness was measured with an optical profiler Wyko NT1100 (Veeco, Tucson, USA) after coating the surface with Au/Pd for better reflection using a sputter coater SC7640 (Quorum Technologies, Lewes, UK). The tests were performed on a home-made tribometer (Fig. 1) that incorporates two units used for driving and measuring purposes. The drive unit consists of a translation stage mounted on a linear bearing and drawn by a dead weight, which is released by an electrical motor at a constant controlled velocity. The measurement unit consists of a 5 kg load cell Z8 (HBM, Darmstadt, Germany) mounted on a hinged balanced arm, which allows loading the contact with known weights and determining the friction force arising due to the motion of the specimen relative to the ball"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003185_012049-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003185_012049-Figure1-1.png",
        "caption": "Figure 1. Maxwell model of induction motor.",
        "texts": [
            " Because FEM provides more accurate information during fault condition contrasted with the expository technique which utilizes direct properties[3]. FEM includes non-linearity such as magnetic saturation, actual properties of magnetic materials, windings placed in actual slots etc., FEM is used for analyzing IM faults such as stator winding faults [7] [8] [9], broken rotor bar faults [10] [11] and bearing faults [12]. The investigation of eccentricity flaw is completed on a three-phase IM with the help of ANSYS Maxwell and Simplorer. The motor is taken care of from a three-phase PWM inverter. Figure 1 shows the IM model in ANSYS Maxwell FEA instrument, the structure subtleties of which are appeared in the Appendix. Static eccentricity fault is incorporated in the IM model in ANSYS Maxwell for a fault degree of 0.15mm (43%). Figure 2 delineates the ANSYS Simplorer model of SPWM controlled inverter fed ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012049 IOP Publishing doi:10.1088/1757-899X/872/1/012049 IM. The IM model in ANSYS Maxwell is simultaneously simulated with Simplorer for winding excitation and load arrangements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002431_sta.2013.6783097-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002431_sta.2013.6783097-Figure5-1.png",
        "caption": "Fig 5. Direction of the displacement of the time-varying midpoint",
        "texts": [
            " Where: M z t ZZ Z , M z t ZZ Z (19) z \u03bcz \u03bcz1(t1 ) \u03bcz1(t0 ) \u03b1(t0) \u03b1(t1 ) z1(t ) z2(t) 1 \u03bcz2(t1 ) \u03bcz2(t0 ) \u0394\u03bcz2 \u0394\u03bcz1 The value of \u03b1 is fixed as the midpoint of the universe. Fig 4. Membership functions with \u03b1 fix And we propose to calculate the premise variable z t by the following: zi t M1i zi t . \u03b1 M2i zi t . Zi M z t . Z M z t . \u03b1 (20) Where: \u03b1 \u03b1,\u03b1 \u2282 Z , Z (21) M z t ZZ , M z t Z , M z t Z , M z t Z Z (22) M1i, M2i, M3i and M4i are the membership grades of z t with \u03b1 time-varying. Let \u03b1R(t) the right displacement of the midpoint on the universe, ensured by a function depending on the error as: figure 5 \u03b1R t fR e t \u03b8 R 1 e \u03b8 R.| | (23) Where: \u03b8 R,\u03b8 R : are respectively the maximum and the grow rate of the \u03b1 R(t). \u03b1L(t) the left displacement of the midpoint on the universe, ensured by: \u03b1L t fL e t \u03b8 L 1 e \u03b8 L.| | (24) Where: \u03b8 L,\u03b8 L : are respectively the minimum, the decay rate of the \u03b1 L(t). The direction criterion depends on the relation between the membership grades and the control law based on the error distance e(t),see subsection 3.1. In this note we use a PDC controller and we based on the relationship (10); the displacement of the \u03b1(t) midpoint to both left or right is directed by the position of the premise variable Z(t); i",
            " If the premise variable is set to the left of the midpoint then \u03b1(t) must approach to the minimum \u03b1 by the function defined in relationship (24). If the premise variable is set to the right of the midpoint then \u03b1(t) must approach to the maximum \u03b1 by the function defined in relationship (23); where the relationship (21) is checked. So, we propose that the displacements \u03b1(t) will follow the Z(t) premise variable by an acceleration determined by the output reference of the Mamdani fuzzy model, figures 6-7. The switch function between the right and the left displacement is illustrated in figure 5 and is given by this sub-program: If \u03b1 t 1 t then \u03b1 t \u03b1R t (25) Else \u03b1 t \u03b1L t It is simple to demonstrate that equations (18) and (20) are exactly equal whatever the value of \u03b1(t); by substituting respectively equation (19) in (18) and (22) in (20).This implies that the subsystems [Ai Bi; Ci 0] of T-S [2,8] does not change suchlike the value of \u03b1(t), and also the criteria of the stability (Stability theorems existing [27-30]) does not change . E. Identification of \u03b1(t) parameters: In this part we use an online identification of \u03b1(t) function shown by figure 6 based on the traditional MRAC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003994_978-3-319-10924-4_2-Figure2.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003994_978-3-319-10924-4_2-Figure2.2-1.png",
        "caption": "Fig. 2.2 The ATRON self-reconfigurable robot is performing one self-reconfiguration step. a-b First, the top right, dark module rotates the white module to its new position. c-d Once at the new position the white module extends connectors to attach to the new neighbour module as can be seen in the bottom right photo between the dark and white module",
        "texts": [
            " typically provide enough computation power to be able to run embedded variants of Linux. \u2022 Batteries are typically Li-Ion allowing modules to functioning for an hour or two. \u2022 Communication may be based on infrared communication, Bluetooth, or WiFi communication. In some cases electrical contact is made between neighbour modules allowing the modules to communicate across a shared CAN bus or similar technology. In a typical self-reconfigurable robot, a self-reconfiguration step consists of a series of small steps as illustrated in Fig. 2.2. First a module disconnects from some of its neighbours, it then moves to a neighbour lattice position, and, finally, extends connectors to attach to neighbour modules at the new position. The mechatronics design of self-reconfigurable robots is a major challenge given the physical constraints and the high requirements in terms of functionality. However, mechatronics is not the focus of this chapter so the interested reader can find more information on this topic in [10, 18]. As we have already argued lattice automata have features that match the desired features of a controller for self-reconfigurable robots"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003035_012007-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003035_012007-Figure1-1.png",
        "caption": "Figure 1. The structure of PRSM.",
        "texts": [
            " In this paper, the vibration data of the planetary roller screw mechanism in two states with and without grease are collected, and features are extracted from the time domain, frequency domain and timefrequency domain, respectively. The predicted accuracy of SVM and BSA-SVM is compared, and the feasibility of the proposed method is verified. Planetary roller screw mechanism (PRSM) is a very important transmission element in precision drive system, which can realize the mutual transformation between rotary motion and linear motion based on the characteristics of large thrust [1], high-precision [2] and higher speeds [3], and is widely used in precision machine tool [4], robots [5], medical equipment [6] and other fields. Figure 1 shows the structure of PRSM. Fault diagnosis, which can detect anomaly and identify failure mode timely, is a very important part to ensure safety and reliability of machinery equipment in condition-based maintenance (CBM) system [7]. In recent years, due to the significant growth of monitoring data and the rapid development of machine-learning technology, data-driven diagnosis methods have been widely developed in bearings [8], gears [9] and ball screw mechanism [10]. For the PRSM, fault diagnosis is the main mean to monitor its safe operation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001145_chicc.2014.6896609-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001145_chicc.2014.6896609-Figure2-1.png",
        "caption": "Fig. 2 The prototype UH with coaxial rotor and ducted fan configuration",
        "texts": [
            " The prototype UH, with net weight 180Kg and height 1.9m, is a vertical takeoff and landing aircraft which includes coaxial rotor with a fuselage with toroidal portion, as shown in Fig. 1. A duct is formed through the fuselage and extends from the top to the bottom of the fuselage. A propeller assembly is mounted to the top portion of the fuselage with a main rotor, of diameter 4.4m, above the fuselage and a ducted rotor assembly in fuselage compensating the propeller antitorque besides providing some fraction of lift, as shown in Fig. 2. The coaxial rotors, main and ducted, rotate at 800 revolutions per minute in the opposite directions with the main rotor providing about 70% of lift, drag, and pitch and roll movements of UH and the ducted rotor providing about 30% of lift and yaw movement. Two light weight two-cylinder engines, with 79.2 horse power, are mounted along longitudinal axis of symmetry below ducted rotor to engage with the rotor assemblies providing sufficient thrust force to lift the prototype UH, as shown in Fig",
            " In order to ensure takeoff and landing safety, UH uses inertia-radio device to measure the height, which makes full use of the radio altimeter\u2019s high precision at low altitudes. In comparison with the main rotor and tail rotor configuration, the prototype UH with coaxial rotor and ducted fan configuration can provide more overall lift and move easily toward an arbitrary direction. This means that these design features not only increase the maneuverability of UH but also increase its stability making it easier to fly during this process. The reference coordinate system of UH is located at its centre of gravity as shown in Fig. 2. X-axis is roll axis, pointing forwards along the symmetry axis. Y-axis is pitch axis, pointing outwards. Z-axis is yaw axis, pointing upwards. The linearized equations of motion for the full 6-DOF can be written as: x A x B u (1) where [ , , , , , , , , ]x y z x y zx and 7 , , ,a e Tu represent the motion states and controls, respectively, of which , ,x y zv v v are forward, lateral and vertical velocity; , ,x y z are roll, pitch and yaw rate; , , are roll, pitch and heading angles; 7 , , ,a e T are main rotor collective pitch, lateral cycle pitch, longitudinal cycle pitch and fan collective pitch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure12-1.png",
        "caption": "Figure 12. Two types of rectangular Bennett linkages: (a) Type A(\u00fe)\u00bcType B(\u00fe) and (b) Type A( )\u00bcType B( ).",
        "texts": [
            " Some insight into the distinct conformations of the two kinds A and B of Bennett isogram is given by noticing that, for a given length of the (BiBj, i\u00bc 1,3; j\u00bc 2,4) diagonal in both kinds A and B, the lengths of the (OM) diagonal in kinds A and B are generally different. Moreover, when the Bennett configuration is rectangular, the two RR bars with a right twist angle are achiral, i.e. cannot be discriminated from their mirrored images. The four types of Bennett linkages fall down into only two types because \u00bc /2 and \u00bc /2 characterize two RR bars that are congruent. Then, the type A(\u00fe) is the same as the type B(\u00fe) and the type A( ) is equivalent to B( ). Both types are shown in Figure 12. The distinction of four types of Bennett isograms is not fully valid for equilateral isograms (a\u00bc b), which can be associated with horn OCTs (Figure 6). An equilateral isogram is an achiral object and there are only two significant types of equilateral Bennett linkages, which are shown in Figure 13. In the kind A, the four bars of a loop are congruent. Type A(\u00fe) is equivalent to Type A( ); only the handedness of four congruent bars differs in A(\u00fe) and A( ) types. Such a loop is degenerated and has partitioned mobility with two one-DoF modes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000252_iros40897.2019.8967673-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000252_iros40897.2019.8967673-Figure2-1.png",
        "caption": "Fig. 2: a) Two-link biped model. Relative hip angle q1 is actuated; absolute stance angle q2 is unactuated. b) Fivelink biped model with curved feet. Relative hip (q1, q2) and knee (q3, q4) angles are actuated; absolute torso angle q5 is unactuated. Parameters in Table I match ERNIE robot [15].",
        "texts": [
            "00 \u00a92019 IEEE 2279 used to orchestrate stable gait transitions on both models. Section IV reports simulation studies to assess the validity of the switching logic. For walkers underactuated by one DoF, the reduced RoA can be represented by a 2D surface. For 3D walkers with two degrees of underactuation, the reduced RoA becomes a 4D surface, which can be estimated offline via numerical computations. Thus, the approach is not limited to planar bipeds, but is applicable to 3D walkers as well. The two- and five-link planar biped models in Fig. 2 consist of rigid links and are left-right symmetric, so one can search for periodic walking gaits using a single step. Bipedal walking dynamics can be modeled as a series of finite-time single support phases and infinitesimally short ground impacts. For a planar biped with n rigid links and m actuators, generalized coordinates q = [q1, ..., qn]T \u2208Q can be used to derive the equations of motion during single support. M(q)q\u0308 + C(q\u0307,q)q + G(q) = B(q)u, (1) where Q is an open subset of [0, 2\u03c0)n that represents feasible configurations of the robot, M(q) \u2208 Rn\u00d7n is the inertia matrix, C(q, q\u0307) \u2208 Rn\u00d7n is the Coriolis matrix, G(q) \u2208 Rn is the gravity vector, B(q) \u2208 Rn\u00d7m is the input matrix, and u \u2208 U \u2286 Rm is a vector of independent control inputs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002174_amc.2016.7496397-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002174_amc.2016.7496397-Figure3-1.png",
        "caption": "Fig. 3. Linkage model of a human body.",
        "texts": [
            " i is the identification character (for example, in the case of the knee joint, i is k): 2) Derivation of the Required Traction: The assistance system is designed in such a way that patients lean on a pad and grasp an armrest while standing with our assistance (see Fig. 1(a)), which means that our system uses the pad to apply force to the patient\u2019s chest and the armrest to apply force to their forearm. These forces move vertically (at the pad) and horizontally (at the armrest). Considering these conditions, we use a linkage model that approximates the human body with our assistance device (see Fig. 3). This model consists of six linkages. The armrest applies the assistance force ( armrestf ) to the center position of Link 1 and the support pad applies the force ( padf ) to the center position of Link 3. im is the mass of the link ( 6,,1i ) and iI is the moment of inertia.  ii yx , is the position of the center of gravity on each link, and  ii yx , (i = a, k, w, s, and e) is the position of each joint. We assume that each linkage is in pillar form with its mass distributed uniformly:                           req swsws wsws ww ypadwxpadw yarmresttsxarmrestts req w Igyxxxyym gyxxxyym gyxxxyym fxxfyy fxxfyy        33111 222 33333 33    (2)                               req wwkwk wkwk wkwk kk ypadyarmrestkw xpadxarmrestkw req k Igyxxxyym gyxxxyym gyxxxyym gyxxxyym ffxx ffyy         44111 222 333 44444     (3) 3) Derivation of the Maximum Traction muscles can generate: We know from previous research [13] that the maximum force that each muscle can realize at the ankle joint is 1meF , 2meF , 3meF , 1mfF , 2mfF , and 3mfF , and the output distribution of the force at the ankle joint is expressed kinematically as a hexagon (see Fig",
            " We know from previous research [8] that the human body motion consists of a voluntary movement which mainly generates the body motion and a postural adjustment which keeps the body stability during the motion. This means the robot should assist with the force when the physical activity is large by a voluntary movement, and should help to adjust the body balance according to a postural adjustment of the patient. Therefore, we propose the following idea;  The first, we define the body movement vector P as (9). This shows the moving direction of the position P (see Fig. 7) which supports the patient\u2019s body (at the center position of the trunk (Link 3), see Fig. 3). The position of P   pp yx , is control reference based on the recommended standing motion by the nursing specialists (see Fig. 6(a)).           ssysyssxsxt pppp \u02c6\u02c6\u02c6,\u02c6\u02c6\u02c6 P (9)  The second, we assume the subject applies all forces userf to the position of P because the support pad and the armrest of our assistive system are connected rigidly. We can calculate userf by the applied force to the pad padf and the armrest armrestf using force sensors in the robot\u2019s body as (10). armrestpaduser fff  (10)  At last, we divide the applied force userf into the parallel force vmf to the body movement vector and the vertical force paf (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003115_physrevfluids.5.053103-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003115_physrevfluids.5.053103-Figure2-1.png",
        "caption": "FIG. 2. Schematic figure. (a) N cilia uniformly distributed at the inner surface of the stationary TaylorCouette device. Np particles are freely suspended in the fluid domain bounded by {(x, y)|R2 2 < x2 + y2 < R2 1}. (b) The snapshots of the beating pattern extracted from [19]. Color coded by its phase \u03c4 .",
        "texts": [
            " A fourth-order explicit Runge-Kutta method is used for evolving both the cilia and the rigid particles. The current capabilities of this hybrid method are demonstrated by simulating the cilia-driven flow within the planar cross section of a fallopian tube as shown in Fig. 1. Note that while the method can handle such arbitrary shapes, for simplicity and ease of analysis, the results section in this paper only considers flows in relatively classical geometries [e.g., Taylor-Couette channel; see Fig. 2(a)]. We emphasize here that such geometries are still difficult to handle using existing methods such as MRS with image systems. The paper is organized as follows. We give the problem formulation and describe our numerical solvers in Sec. II. An analysis of the mixing and transport properties of actuated cilia in complex domains will be presented in Sec. III, followed by conclusions and future work in Sec. IV. In this section, we first describe the problem formulation for the specific case of cilia-driven flow of rigid particles suspended in a Taylor-Couette device",
            " Then, we show how to recast it as a set of mixed boundary integral and discrete equations with unknowns residing on the cilia, particle, and wall boundaries only (thus leading to dimensionality reduction). Lastly, we describe a numerical method for discretizing and solving these equations. 053103-3 Consider a thin gap of fluid confined between two stationary concentric circles of radius R1 and R2 with R1 > R2. The fluid domain is denoted by = {(x, y) | R2 2 < x2 + y2 < R2 1}. Np rigid particles are immersed in the fluid; N cilia of length are uniformly distributed on the surface of the inner circle [see Fig. 2(a)]. Following [19], the kinematics \u03b6 of each cilium in its body frame can be approximated by a truncated Fourier series in time \u03c4 and a Taylor series in arclength s. The resulting beating pattern is shown in Fig. 2(b). We apply proper rotations and translations to \u03b6 to obtain the position of the lth cilium r such that the cilia are uniformly distributed along the inner circle and are oriented perpendicular to the circle. Specifically, the position of the lth cilium at arclength s and time t is given by rl (s, t ) = [ cos(\u03b8l ) sin(\u03b8l ) \u2212 sin(\u03b8l ) cos(\u03b8l ) ] \u03b6(s, \u03c4l ) + R2 ( sin(\u03b8l ) cos(\u03b8l ) ) , \u03b8l = 2\u03c0 (l \u2212 1)/N, \u03c4l = 2\u03c0t + (l \u2212 1) \u03c6. (1) By construction, the first cilium is rooted at (x, y) = (0, R2) and the index of the cilium increases clockwise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002745_s12206-019-1235-8-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002745_s12206-019-1235-8-Figure6-1.png",
        "caption": "Fig. 6. Impedance controller with amnesia removal.",
        "texts": [
            " This error developed is because of the decrease in the gain of overwhelming controller (\u03b1) for the duration of interaction period such that force control is attained at the expense of correctness of trajectory tracking. This position error built up at the end of a step of force control is known as amnesia. This amnesia error must be eliminated before the next interaction so that it follows the reference trajectory during the period when it is not interacting. The impedance controller with removal of amnesia is represented in the form of a block diagram as shown in Fig. 6. The force controller builds up the loss in positional information ( )pd during the period of interaction. This information of position is calculated by integration of signal of flow ( )pd& that is proportional to the error of force i.e. the flow in bond number 3 in the bond graph arrangement shown in Fig. 5(a). Another flow variable ( )red is given in the inverse system with velocity of reference (Vr) for the duration of period when it is not interacting. This flow variable is proportional to the amnesia error",
            " This amnesia removal rate ( )fd is a term derived from trial and error [29, 30] and is given as ( ) ( )0 f en lim x dtt tG e F Fd V-= -\u00f2& (4) where G is a gain and 0t is the initial instant, i.e. the point at the end of a step of force control. Therefore, in addition to the velocity of reference, this signal of flow ( )fd& is also given to the inverse controller. The quantity of accommodated position recovered is also noted down simultaneously such that ( )p p f dtd d d= -\u00f2 & & . The function ( )f .j in Fig. 6 models the Eq. (4). The various parameter values applied for the simulation are shown in Table 1. The values of parameters for plant, pad and inverse controller for the hybrid manipulator have been chosen suitably according to the application. The gain \u2018\u03b1\u2019 of overwhelming controller is intentionally taken as 100 (i.e. >>>1) here and the value of proportional controller gain \u2018 \u2019z is taken near to zero because a high value of gain implies impedance control strategy in a better way and there is minimum error in trajectory tracking, as already explained in above section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.17-1.png",
        "caption": "Figure 2.17 Limb electrodes: (A) curved plate, (B) fixation elastic bands and (C) plate electrode mounted in clip support.",
        "texts": [
            " To avoid chemical reactions with electrolyte or sweat, that may cause skin irritation, the electrodes are made, preferably, of inert or almost inert metals and metal alloys such as gold, platinum, silver, titanium, and stainless steel. The metal electrodes class can be classified as limb, suction, disc, EEG, floating (top-hat and disposable), and dry electrodes. This type of electrode usually has rectangular format and plane contact area. Its surface can also be curved to a better adjustment to the format of arms and legs (Figure 2.17A). They are used, for instance, as return plate in electrosurgery and right leg electrode in ECG recording. The inner side of the electrode receives a layer of electrolytic gel and it is fixed with the help of elastic bands. Smaller sizes of plates are also mounted in clip supports (Figure 2.17B), eliminating the need of the bands. In addition to gold, silver, and stainless steel, they can be made of a metallic alloy of copper, zinc, and nickel, known as German silver or alpaca. German silver is extensively used because of its hardness, toughness, resistance to oxidation, and high electrical resistance; the percentage of the three elements varies, ranging for copper from 50% to 61.6%; for zinc from 17.2% to 19%; and for nickel from 21.1% to 30%. The suction electrode has a metallic dome shape with a connector to the electric cable, and a suction rubber bulb (Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002481_00207543.2016.1259669-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002481_00207543.2016.1259669-Figure1-1.png",
        "caption": "Figure 1. The geometry model based on \u2018body-flange\u2019 structure.",
        "texts": [
            " Based on the above analysis, we organise the rest of the paper as follows. In Section 2, we introduce a parametric CAE modelling method for HPTD. Section 3 trains a BPNN with virtual samples. Section 4 proposes a hybrid initial population for non-dominated sorting genetic algorithm-II (NSGA-II). In Section 5, in order to validate the feasibility and efficiency, the proposed framework is applied to optimise an HPTD. Section 6 gives some conclusions and future works. HPTD is a product with complex geometrical details, as Figure 1(a) shows. It is difficult to make a full analysis due to its complicated working conditions. Considering the feature of axial symmetry, two-dimensional FEM can be used to reduce the cost of CAE. Furthermore, to meet the optimisation requirements such as the speed of samples generation and convenience of parameter management, a template is used for parameterisation. 2.1 \u2018Body-flange\u2019-based parametric geometry modelling A turbine disc consists of body, flanges and mortises. Among them, a body consists of the rim, hub and web plate that are connected by different types of profiles; flanges connect the disc with other components; mortises connect the blades and disc, and information of mortises is fused into the rim to make an equivalent \u2018body-flange\u2019 model for twodimensional FEM. Figure 1(b) and (c) describe the geometric details of a typical HPTD. Body parameters include rim radius R1, rim thickness W1 and rim height H1, shaft radius R2, hub thickness W2 and hub height H2, web plate\u2019s outer diameter R3 and inner diameter R4, outer thickness W3 and inner thickness W4; flange parameters are the location (H3, W5, H5, W7), length (H4, H6\u2013H5), width (W6, W8) and chamfer (R5, R6). A parametric modelling method is used to support the extensibility of the aforementioned \u2018body-flange\u2019 structure, which is shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002404_mmar.2016.7575286-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002404_mmar.2016.7575286-Figure6-1.png",
        "caption": "Fig. 6. The frames of: robot station, robot base, object, working link, and gripper",
        "texts": [
            " THE ROBOT DESIRED TRAJECTORY To general description the robot desired trajectory we apply the frame xsyszs, associated with the robot station [8]. As mentioned in Section 2 reference frame xyz is also associated with the robot station. We assume that frames xyz and xsyszs are identical. The robot desired trajectory describes the movement of the gripper to the object observed by the cameras. With the gripper is associated a frame x7y7z7 [8]. Matrix obT7d(s) describes the frame x7y7z7 relative to the frame xobyobzob. The parameter s is the length of the gripper movement path calculated from the starting point of the trajectory. Figure 6 illustrates the following frames: robot station xsyszs, object xobyobzob, robot base x0y0z0, working link x6y6z6 and gripper x7y7z7. From the figure results (6) describing the matrix T6d(s). 1 76 )()( \u2212= ETTTT ss d ob obsd . (6) The matrix T6d(s) describes the desired trajectory of the frame x6y6z6 relative to the frame x0y0z0. The user of robot defines the movement of the gripper relative to the manipulation object using matrix obT7d(s). The most convenient define the beginning of the trajectory at a point above the manipulation object",
            " 7 in the frame xpypzp. From Fig.4 results the frame xpypzp as in Fig. 9, described by (12).             \u2212 = 1000 0100 5.67010 5.42001 ob p T . (12)             \u2212\u2212 \u2212\u2212 \u2212\u2212\u2212 = 1000 3154.59999.00063.00014.0 0780.10063.09999.00140.0 9511.00015.00140.09999.0 obT . (13)             \u2212 = 1000 0100 0001 388010 sT . (14)             \u2212\u2212 = 1000 45010 146100 0001 E . (15) The matrix Tob (see Fig. 5) calculated from (5) has the form (13). For the robot station on which the studies were carried matrix Ts (see Fig.6) has the form (14). The matrix E describing the gripper frame x7y7z7 relative to the working link frame x6y6z6 has the form (15). It was assumed that the desired trajectory of the gripper relative to the object is a straight line coinciding with the axis zob. The gripper has to start straight motion from the point S and end at the point F above the object. It was assumed that the point S is at a height of 314 mm above the plane xobyob as in Fig.10. For such defined trajectory of the gripper we obtain the matrix obT7d(s) described by (16)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002092_j.mechmachtheory.2016.04.003-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002092_j.mechmachtheory.2016.04.003-Figure1-1.png",
        "caption": "Fig. 1. Force analysis of the complete leg.",
        "texts": [
            " All rights reserved. In paper [1,2], considering the rotational equilibrium of the lower part and upper part of the leg in the fixed leg frame, Euler's equation gives \u2212mdrd ad \u00femdrd g\u2212IdA\u2212W IdW \u00feMus\u2212r Fp\u2212Mp\u2212CuW \u00bc 0 \u00f01\u00de \u2212muru au \u00femuru g\u2212IuA\u2212W IuW \u00fe S Fs \u00fe r Fp \u00feMp\u2212Cs W\u2212\u03c9\u00f0 \u00de \u00bc 0 \u00f02\u00de where Fp is the vector force at the prismatic joint exerted by the lower part on the upper part acting at a point r, Mp is the vector moment at the prismatic joint acting on the upper part, other parameters refer to Fig. 1. The basic error of these equations is that the angular momentum theorem isn't properly applied. Let O0 be a point fixed in inertial space, C is the center of mass and B is an arbitrary point fixed on the body (see Fig. 2). In combination with the vector operation rules and Newton's axioms, the angular momentum theorem for a rigid body is obtained in the final form m\u03c1c \u20acrB \u00fe JB _\u03c9\u00fe\u03c9 JB \u03c9 \u00bc MB \u00f03\u00de 3 in which \u03c9 is the absolute angular velocity of the body, \u20acrB, JB and MB are the absolute acceleration, inertia tensor and resultant moment about point B respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003978_978-3-319-14705-5_12-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003978_978-3-319-14705-5_12-Figure9-1.png",
        "caption": "Fig. 9 Straight motion",
        "texts": [
            " (15) the dynamic model of spherical robot BHQ-1 can be deduced as \u23a7\u23aa\u23a8 \u23aa\u23a9 7 5 mr2\u03c9\u03071 = m0 1 \u2212 r f 0 2 7 5 mr2\u03c9\u03072 = m0 2 + r f 0 1 2 5 mr2\u03c9\u03073 = m0 3 (23) where, m0 1, m0 2, m0 3 are the projections of the principal moment m0 on the three axes of the body reference frame {Ob X \u2032Y \u2032Z \u2032}, f 0 1 , f 0 2 are the projections of the principal force f 0 on axes X \u2032, Y \u2032 of frame {Ob X \u2032Y \u2032Z \u2032}, f 0 and m0 are the principal force and principal moment imposed on the geometric center of spherical robot BHQ-1 respectively. In Fig. 9, frame {o\u2032ijk} is located on the geometric center of BHQ-1 and its orientation is the same as that of frame {oxyz}. When spherical robot BHQ-1 moves along a linear trajectory its hollow axle and those installed components will rotate around axis i to reach a high position supposed as the one shown in Fig. 9. Because there is no rotation about axes j and k, \u03c8 = 0 and \u03d5 = 0 are obtained, substituting them for the variables in Eq. (16), we can get those quasi-velocities as \u03c91 = \u03b8\u0307 , \u03c92 = 0, \u03c93 = 0, \u03c94 = 0, \u03c95 = 0 (24) In Fig. 9, the gravity direction is along the \u2212k direction (downward vertically), so the projection of gravity on plane io\u2032 j is zero, that it to say f 0 1 = 0, f 0 2 = 0, substituting them for the variables in the dynamic model (23), the following simplified dynamic model of BHQ-1 can be got. \u23a7\u23aa\u23a8 \u23aa\u23a9 m0 1 = 7 5 mr2\u03b8\u0308 m0 2 = 0 m0 3 = 0 (25) So the gravity moment exists only around axis i and the mass sways only in the plane o\u2032k j . From Eq. (25) and m0 1 = mg l j + f r (26) we can get the driving moment of motor 1 is M1 (t) = mg l j + mL2\u03b2\u0308 = \u22127 5 mr y\u0308(t) \u2212 f \u00b7 r + mL2\u03b2\u0308(t) (27) where, L is the distance between the center of the mass and the rotation axis of the hollow axle, \u21c0 l j is the projection vector of L on axis j, \u03b2 is the angle that the mass deviates from axis k, f is the friction vector imposed on the spherical robot by the ground, g is the gravity acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001850_978-3-319-23335-2_2-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001850_978-3-319-23335-2_2-Figure2-1.png",
        "caption": "Fig. 2 Different stages of the construction of the wing sail prototype, from the CAD drawings to the final result",
        "texts": [
            " We\u2019ve decided to limit the length of the sail to one meter and the cord to 40 cm, yielding a design that we could easily transport, if detachable from the main structure. We\u2019ve also analyzed the impact of the aspect ratio in construction and control and we\u2019ve decided to use a NACA 0015 profile, a well studied design with a symmetrical shape enabling similar lift generation in both sides of relative orientation. The first sail prototype was assembled from a core of balsa wood with aluminum and epoxy reinforcements and wrapped in a thermal foil. Figure 2 shows some of the construction stages. To ensure the overall NACA 0015 profile, 10 cross sections were machined from 5 mm thick balsa wood, to become the backbone of the sail. These were drilled with guiding holes to ensure a proper alignment: a central hole with 20 mm diameter for an aluminum tube that is used as the sail mast, and smaller holes for 8 mm aluminium tubes that reinforce the structure. To facilitate construction and avoid sharp edges, we\u2019ve modified the leading and trailing edges of the sail to round shapes",
            " These battens were also glued with epoxy before being wrapped with the resin-coated thermal foil, which was carefully glued to the wood using a heatgun. With this assembly, both the top and bottom of the wing sail had sharp edges, which could cause turbulence and affect aerodynamics. To avoid this, we\u2019ve designed two mechanical parts to round both ends of the sail, and we\u2019ve used a 3D printer to fabricate them in ABS plastic. The final assembly weights 1.3 kg and it can be seen in the last picture of Fig. 2. In order to evaluate the aerodynamic characteristics of a single sail under various wind profiles and compare it to the theoretical values of a NACA 0015 profile, we\u2019ve used a Computational Fluid Dynamics (CFD) software, Autodesk Flow Simulation. We\u2019ve conducted simulations for wind velocities of 3, 5, and 10 ms \u22121 , with an angle of attack varying from 0 to 180\u25e6. Figure 3 shows an example of the graphical output of the CFD with the velocity field around the sail. With the resulting forces, we\u2019ve interpolated the lift and drag coefficients in one degree intervals and the results can be seen in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003373_10667857.2020.1810926-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003373_10667857.2020.1810926-Figure6-1.png",
        "caption": "Figure 6. CAD model in CATIA v5 after editing: (a) outside view and (b) inside view.",
        "texts": [
            " Surface is offset by 1 mm to cater the cushion during use of the splint. A thickness of 3 mm is provided to this surface as shown in Figure 5. Two large oval shaped holes and many small holes are created on the CAD model to reduce the weight of the splint as well as to improve the comfort level of patient by promoting aeration through skin-split interface. Finally, a fillet of 1 mm and 0.5 mm radius are created to larger and smaller edges respectively and provision for Velcro strap is designed on the model as shown in Figure 6. The final CAD model is shown in Figure 7. Figure 3. Surface model obtained from mesh in Geomagic Studio. Finite element analysis of CAD model. A finite element analysis (FEA) based simulation is performed on CAD model to evaluate physical behaviour of the splint against unintentional hitting during recovery period. Stress and displacement analysis is carried out in Fusion 360 software (Autodesk, USA) under static loading conditions. Finite Element Analysis is conducted for polyamide-PA2200 material with Young\u2019s Modulus (E) and Poisson Ratio of the order of 1500 MPa and 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.71-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.71-1.png",
        "caption": "Figure 3.71 Belt and pulley schematic",
        "texts": [
            " The side of the belt being pulled by the power pulley is in tight tension (Tt ) while the side of the belt returning from the load pulley is in slack tension (Ts). The effective force supplying power to the load is: Te = Tt \u2212 Ts (3.108) There is no load on the belt or the supply power source other than the belted load and operation is usually continuous. System Components 127 The more recent configuration is the \u201cpositioning drive\u201d in which the belt connects two equally sized pulleys with a load attached to the belt, as shown in Figure 3.71. The use of the belt and pulley instead of a ball screw system is mainly based on the ability of the belt and pulley to achieve higher velocity over longer distances than the ball screw, as shown later in a comparison chart. One pulley is connected to a power source (a servo motor) and the other is simply an idler pulley. Operation is intermittent with the load being driven to various locations between the end limits of the assembly according to a motion profile. The load is typically connected to a ball bearing supported plate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001229_sii.2014.7028114-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001229_sii.2014.7028114-Figure6-1.png",
        "caption": "Fig. 6. Specifications of electrode arrays. (a) is ver.1. (b) is ver.2.",
        "texts": [
            " Proximal phalanx is not suitable for presenting a joint angle because it requires that a center-to-center distance of two electrode is about 7mm or more. In this paper, a type of joint angles is presented on distal phalanx which has a high haptic resolution for electric stimulation on a comparison with other phalanx because at least 9 electrodes which are required to present a type of relative hip joint angles can be mounted on the phalanx if its two-point distance is set at 4mm. IV. IMAGINARY LEGS ANGLE CONTROL EXPERIMENT BY USING THE FINGER-MOUNTED WALK CONTROLLER WITH AND WITHOUT FEEDBACK A. Electrode Arrays Figure 6 shows the electrode arrays for relative hip joint angles presentation. Figure 7 shows movement patterns of current using dotted red lines. An electrode of the electrode arrays has a diameter of about 2mm. Center-to-center distances of two electrodes are 3.59 and 5.5mm. The distances are based on preceding chapter\u2019s results. The whole measurements of these arrays are adjusted for a subject on the experiment. Figure 6 (a) is an electrode array ver.1. The array ver.1 has 10 stimulation points and presents relative hip joint angles every 4.56 degrees by moving current like a dotted red (a) (b) (c) Fig. 5. Answer rates of two-point discrimination experiments every two-point distances. (a) is results of distal phalanx. (b) is results of middle phalanx. (c) is results of proximal phalanx. (a) (b) Fig. 4. (a) The contact surface of small electrode array. It has 9 stimulation points on the center part and 14 grounding points on the peripheral part. (b) An example of a contact with a distal phalanx. 978-1-4799-6944-9/14/$31.00 \u00a92014 IEEE 646 line of Fig. 7 (a). The angle means a maximum angular resolution. Its resolution is the same as the resolution of a device which is used in our previous paper. Figure 6 (b) is an electrode array ver.2. A size of the array ver.2 is almost same as the array ver.1. The array ver.2 has 22 stimulation points and presents relative hip joint angles at every 1.95 degrees by moving current like a dotted red line of Fig. 7 (b). Therefore, the array ver.2 has about 2.34 times larger information capacity than the other. A user covers his/her fingertip of middle finger with a band like Fig. 8 when he/she uses the array ver.1 or ver.2. The band is easy to put on and take off the array because it is fixed by a Velcro and can be stretched"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-11-1.png",
        "caption": "Figure 6-11. Step 4 to obtain the propulsion or allocation matrix",
        "texts": [
            " However, in this example and for a certain degree of standardization we will choose the direction of rotation shown in Figure\u00a06-10. Step 4: Choose the geometric configuration. In this case, place the arms at the same distance to the center of the vehicle, with square angles (0 and 90 degrees) and at the same height with respect to the center of the vehicle. Very unequal distances, very different angles, or significantly different heights change the propulsion matrix considerably. Nevertheless, if the differences are moderate, this does not influence the operation of the vehicle. See Figure\u00a06-11. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 254 Step 5: Place the coordinate frame of the drone (that is the body frame) and label the axes. This influences the vehicle\u2019s movement and control design. The choice will depend on the test of the remote control levers. See Figure\u00a06-12. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 255 Step 6: Relate the rotational frames and their direction of rotation. In this example, you will associate the pitch with the X axis of the drone, the roll with the Y axis, and the yaw with Z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001658_j.biombioe.2013.08.033-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001658_j.biombioe.2013.08.033-Figure1-1.png",
        "caption": "Fig. 1 e Configuration of dual-chamber rumen MFC.",
        "texts": [
            " Rumen contents for the preparation of rumen fluid were collected via the rumen cannula of non-lactational Holstein dairy cattle. The cattle were given a diet of Bermuda grass straw (10 kg) and a commercial concentrate (2 kg), fed in two equal portions at 0800 and 1800 daily. Just prior to themorning feeding, rumen contents were taken and strained through 4 layers of gauze into a pre-warmed (39 C) thermos flask. The resulting rumen fluid was filtered through 8 layers of gauze under a stream of CO2 to further remove small feed particles. MFCs (Fig. 1) were constructed from two plastic chambers (15 cm 14.5 cm 6.6 cm). The chamberswere connected and tightened usingmetal clamps and sealed by silicone. A proton exchange membrane (PEM, Nafion N117, Dupont Co., USA), with a reaction area of 145 cm2 (14.5 cm 10 cm), was placed between the two compartments. The working volume was 1000 cm3 in each chamber. Both anode and cathode electrodes were made of graphite plate and carved into a rectangular shape. The electrode size was 10 cm 10 cm with a thickness of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001364_sii.2013.6776701-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001364_sii.2013.6776701-Figure5-1.png",
        "caption": "Fig. 5 An SMG plate with film heaters.",
        "texts": [
            " The surface of the SMG adheres to the aluminum foil and the aluminum foil slips on the film heater. Figures 8 shows the difference between the SMG with and without aluminum foil. The plate of Fig. 8 (b) becomes softer than that of Fig. 8 (a) for following reason. The film heater is easy to be bent and hard to be extended. The tip of the heater is fixed to the tip of the SMG in Fig. 8 (a) and the tip of the heater is not fixed in Fig. 8 (b). The SMG plate and the heater are bound by rubber bands so that the film heater contacts the aluminum foil as shown in Fig. 5. We select the aluminum foil for following reason. The aluminum foil is thin and easy to be bent, it has low friction and high thermal conductance and its cost is low. Figure 9 shows the temperature control system of the SMG plate. An IC temperature sensor is attached to the SMG. An Arduino board [17] receives the sensor data and transmits the (a) (b) (c) (d) Fig.2 Pulling of a hot SMG. (a) (b) (c) (d) Fig.3 Pulling of a cold SMG. 978-1-4799-2625-1/13/$31.00 \u00a92013 IEEE 648 On/Off command to transistors to drive relays for switching of electric current to film heaters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000767_s12206-013-0108-9-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000767_s12206-013-0108-9-Figure4-1.png",
        "caption": "Fig. 4. Load-carrying analysis diagram of ZT20SW and its coordinate system.",
        "texts": [
            " min max[ , ]b bp p p p p= + \u2206 \u2208 (1) When preload pb comes to pmax and x is still larger than x0\u2019, supporting rigidity kd and cutting force Wt\u2019 on the table needto be analyzed and checked. Load-carrying of NC rotary table can be categorized into two major types: static load and dynamic load. For example, axial force caused by fixture and workpiece on the table is a kind of static load, while discontinued impact load of the table caused by gear milling is a kind of dynamic load. For the convenience of load-carrying analysis, two coordinate systems were established on ZT20SW. The rotating coordinate system O\u2019Z and the linear coordi- nate system OX are shown in Fig. 4. Working plane of ZT20SW rotates about central axis O\u2019Z, while OX is parallel to axis of the worm and tangent to pitch circle of the worm gear. In coordinate system O\u2019Z, load-carrying status of ZT20SW can be described as the following moment equilibrium equation: ' 0( ) 0t e a b d fW R F F F F R M+ + + \u2212 \u00b1 = . (2) In Eq. (2), velocity and acceleration of the table is low while gear milling, so Fe and Fa could be ignored. Cutting force Wt\u2019 is related to its machining parameters of gear milling, which is a kind of high-efficiency manufacturing method and can apply periodic exciting loads on the table"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure18-1.png",
        "caption": "Fig. 18. The typical configurations of plane-motion Mode V.",
        "texts": [
            " 17 , the configuration is also a special configuration of Mode II. Since the mechanism is an open-chain mechanism with 3 links and 2 R joints, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration III ( Fig. 17 ), when the axes of R joints J and J coincide, the axes of R joints J and J coincide, and the axes of R joints J and R2 R4 R6 R8 R1 J R5 coincide during the movement, the mechanism makes a planar motion, which is called plane-motion Mode V. Its typical configurations are shown in Fig. 18 . Due to the characteristics of antiparallelogram unit ( Fig. 19 ), it is a bifurcation point at \u03b8AA = 0 or \u03b8AA = \u03c0 . At a bifurcation point, an antiparallelogram could turn into a parallelogram. Therefore, there are some more special configurations as shown in Fig. 20 (no configurations interfere with each other by the layering). In addition, when interference is not considered, some special configurations are shown in Fig. 21 . Some configurations of plane-symmetric Mode I, plane-symmetric Mode II, plane-motion Mode III, plane-motion Mode IV, and plane-motion Mode V can be transited to each other, and they have a certain relationship as shown in Fig",
            " 10 (c) of plane- symmetric Mode II cannot be achieved because of the offset angle. As shown in Fig. 26 , from the bifurcation configuration in Fig. 26 (a) (which corresponds to configuration (6) in Fig. 22 ), plane-symmetric Mode I ( Fig. 26 (b) and (c)) and plane- symmetric Mode II ( Fig. 26 (d) and (e)) can be obtained. Plane-motion Mode III (corresponding to Fig. 14 ) are shown in Fig. 27 . Figs. 28 and 29 are the snapshots of plane-motion Modes IV (corresponding to Fig. 16 ) and V(corresponding to Fig. 18 ), respectively. Fig. 30 shows the snapshots of the special configurations(corresponding to Fig. 20 ) of this multiple- mode mechanism. Similar to Bricard-like mechanism [61] , the multiple-mode mechanism discussed above can also be used as a construction unit to construct deployable mechanisms. As shown in Fig. 31 , a quadrangular prism deployable mechanism composed of two identical multiple-mode mechanisms connected by eight spherical joints is constructed. During its movement, the two identical multiple-mode mechanisms are mirror-symmetrical"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.23-1.png",
        "caption": "Figure 2.23 Schematic diagram of a reusable metal disc top-hat electrode.",
        "texts": [
            " The floating classification is related to the electrodes contact design made through a thick electrolytic gel layer. Floating electrodes are less susceptible to motion artifact because the metal sensor is recessed, placed inside a casing and do not enter em contact with the skin. The casing fixed to the skin is filled with the electrolyte, in a way that the metal sensor \u201cfloats\u201d in the electrolytic gel and prevents the double layer of charges to be disturbed. This classification includes reusable top-hat and disposable electrodes. 2.6.1.1.5.1 Top-hat electrodes Figure 2.23 shows a schematic diagram of a reusable metal disc top-hat electrode for biopotential recording. The Ag AgCl disc is placed inside a hollow insulating casing and do not touch the skin. The inner part of the casing is filled with electrolyte and a double-sided adhesive tape promotes fixation of the electrode to the skin face. The embedded electrolyte does not move relative to the electrode, assuring that there is no relative movement between the skin and the double layer of charges in the metal/electrolyte interface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002561_iros.2016.7759080-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002561_iros.2016.7759080-Figure4-1.png",
        "caption": "Fig. 4. Open-chains model of each serial hybrid chain.",
        "texts": [
            " IMPULSES AT ORIGIN AND INSERTION POINTS In this section, we summarize the internal impulse model at cut-joints and origin/insertion points of muscle. The external impulse at the distal end and velocity increments at contact point and cut-joint impulses are already known since they were derived in section (II) and (III), respectively. Each serial hybrid chain is transformed to open tree structure to find the internal impulses at origin and insertion point of muscles, by cutting the closed-loop links and joints. The open tree structure consisting of six serial-chains for each serial hybrid chain is shown in Fig. 4. The impulsive forces \u02c6i k cF acts with the same magnitude but opposite in direction on muscle and on link. The general expression for inverse dynamic model for each serial chain (muscle models) for both (left and right) serial hybrid chain is established as follows [ ] [ ] [ ] , ( 1...6)i i T i i c T i k k i k i k k i k k k cI P G k     T F   (18) where [ ]i k I and [ ]i k P denote the inertia matrix and inertial power array of the thk serial-chain in the thi serial hybrid chain, respectively. [ ]i h k G denotes the first-order KIC relating the cut joint velocity to joints velocity ( k  ) of th k serial-chain in each thi serial hybrid chain"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001575_1464419314540901-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001575_1464419314540901-Figure2-1.png",
        "caption": "Figure 2. Shaft finite element model.",
        "texts": [
            " Then, two types of resonance for the VC excitation in low speed region are discussed mainly with the results of cascade plots and frequency spectrums. A typical roller bearings\u2013rotor system with local nonlinearity is shown in Figure 1, which can be divided into linear component (the elastic shaft without bearings) and local nonlinear components (bearings and disk). The system under study has the outer race of the bearing fixed to a rigid support and the inner race fixed rigidly to the shaft. A 2-node Timoshenko element model with 8 degrees-offreedom (DOF),19 as shown in Figure 2, which can account for the effect of inertia moment, and shear deformation, is employed. With the help of Hamilton principle and finite element, the systematic dynamic equation can be described as M \u20acq\u00feG _q\u00fe K q \u00bc Fb \u00fe Fu \u00fe Fg \u00f01\u00de where M and G respectively stand for the mass matrix and gyroscopic matrix of the rotor after considering the disk, K is the stiffness matrix of the rotor, Fb is the vector of nonlinear supporting forces of bearings, Fu is the periodic exciting force (i.e. unbalanced excitation with the same phase as the rotating speed) vector acting on the rotor, and Fg is the gravity vector of the rotor, and q\u00bc (x, y, , \u2019) is the displacement vector of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure7-1.png",
        "caption": "Figure 7. Velocity vector contours for standard dimensions in the X direction.",
        "texts": [],
        "surrounding_texts": [
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nincrease in the flow rate is observed, due to the narrowing of the space between the drum and the housing wall.\nA slight discrepancy in the display of velocities in the graphs of displaying fields and vectors is due\nto the peculiarities of displaying and visualizing numerical results.",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nAn optimization calculation was carried out to assess the effect of the hull geometry. The original body geometry (see Figure 1) has been changed by moving the top down 70 mm - Figure 9.\nFigures 10 - 12 show the contours of the velocity fields and the velocity vectors for the modified housing shown in Figure 9. The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 31.2 m/s.\nFor this geometry, a more even distribution of the air velocity fields is observed."
        ]
    },
    {
        "image_filename": "designv11_30_0000025_rnc.4734-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000025_rnc.4734-Figure3-1.png",
        "caption": "FIGURE 3 Fin configuration",
        "texts": [
            " For majority of interceptions, T\ud835\udefc is considered as a known slowly varying quantity and can be discussed as following equation: T\ud835\udefc = \ud835\udefc . \ud835\udefeM = V ( m \ud835\udf0cS )( \ud835\udf15cl \ud835\udf15\ud835\udefc )\u22121 , (2) where \ud835\udf0c = \ud835\udf0cV 2 M\u22152 denotes the kinetic pressure. In this paper, T\ud835\udefc is considered as a constant. Remark 2. The relationship among the angles \ud835\udf06, \ud835\udefeM, \ud835\udf03, and \ud835\udefc is shown as Figure 1. The angle of attack \ud835\udefc can be calculated via (1), and other angles can be measured by inertia sensors on the controlled missile. Remark 3. Assume that 4 fins are configurated as \u201c+\u201d structure, as shown in Figure 3, and then, fin deflection can be calculated by \ud835\udeff = (\ud835\udeff1 \u2212 \ud835\udeff3)/2. Because interceptor is an axisymmetric structure, fin deflection \ud835\udeff is equal to \ud835\udeff1 and \ud835\udeff3 numerically. Assuming that both the interceptor and the target are point masses, the planar homing engagement dynamics is depicted as in Figure 2, where M represents the interceptor and T represents the target, and other denotations can be found in the nomenclature. According to Figure 2, some equations can be concluded to describe the interception geometry at the end game of guidance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003672_icem49940.2020.9270785-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003672_icem49940.2020.9270785-Figure1-1.png",
        "caption": "Fig. 1: a) FEA source mesh S, b) congruent target mesh T of the rotor core.",
        "texts": [
            " A second FE mesh V(T (\u2126),R2) \u2208 V(S,R2) which is congruent with the original mesh of the least symmetrical model part is created. The local magnetizations MT \u2208 V(T (\u2126),R2) for the elements eTi of the target mesh can be calculated by MT i = Ne,S j\u2211 k=1 kTj \u00b7M S k \u00b7 T\u03a8 . (16) The factor kTj = dSe,Sj Se,Sj (17) represents the local scaling of MS according to the overlapping elements eTj and eSi when the source mesh is moved time-step wise below the target mesh and thus, corresponds to a local weighting of the magnetization between the source and the target mesh. Fig. 1 shows the a) source mesh S respectively the congruent target mesh T of the rotor core region of an example 3-phase, 4 pole PMSM. To calculate the local scaling factors corresponding to (17) in a first step a virtual 360 degree source mesh is created according to (14) and (15), shown in Fig. 2 a). Subsequently, for each element of the target mesh T the overlapping elements from the source mesh S are identified. In Fig. 2 b) one target mesh element is depicted in red, the relevant source mesh elements respectively in green"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001743_iccas.2015.7364719-FigureI-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001743_iccas.2015.7364719-FigureI-1.png",
        "caption": "Fig. I Appearance of the HAUV under development.",
        "texts": [],
        "surrounding_texts": [
            "Recently, there have been an increasing interest on marine resources, and also there have been a growing need for marine equipment. Underwater robot among various marine equipment is essential equipment for ocean exploration and resource development. Thus underwater robot field is now growing rapidly due to the exploration, the investigation, and the research activities of deep sea [1-4]. AUV(autonomous underwater vehicle), a kind of the underwater robot, is able to cruise at high speed in the sea using its own power supply without any power line or signal wire, and can perform its mission [4]. AUV is classified into two groups, crulsmg AUV and hovering AUV. HAUV(hovering AUV) is the title for the underwater robot which conducts detailed investigation and works of specific area. HAUV must have the function of attitude and position control on the spot and the function of way-point movement in order to perform detailed examination and works [5]. PID controller and fuzzy controller are utilized for a certain level of control when there isn't a dynamical model about that underwater robot and its environment [6-8]. In this paper, way-point chasing simulation using PID controller as Matlab/Simulink is covered for tracking HAUV way-points and attitude control which are now being developed in our lab like Fig. 1."
        ]
    },
    {
        "image_filename": "designv11_30_0001227_ssrr.2013.6719358-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001227_ssrr.2013.6719358-Figure2-1.png",
        "caption": "Fig. 2: NIFTi robot; All measurements are in millimeters",
        "texts": [
            " Then, the column map C(i, j, t) is inductively defined as follows: C(i, j, t) =P(t), for t=0 C(i, j, t) =C(i, j, t\u22121)\u222a {p=(x, y, z)>\u2208P\u2212(t) |(i, j)= ( sgn(x) \u2308 |x| r \u2309 , sgn(y) \u2308 |y| r \u2309) } for t > 0 (2) Here |x| indicates the absolute value of x, sgn(x) the sign, and dxe is the ceiling operator, mapping x to the smallest following integer. Note that the point cloud is continuously updated and new detected points at time t are mapped to the appropriate column maps. The set of points falling into the column map C(i, j, t) is further partitioned into a set of levels. Each level is an interval of the form [Zmin, Zmax] obtained by partitioning the z according to height parameter RH (see Figure 2). Definition 2 (Levels set): Let a level be defined as: `=[Zmin, Zmax] (3) let p=(x, y, z)>\u2208P(t), \\ be the set difference operator, and let us abbreviate min and max with M= max, m= min, just in this definition for sake of space. Then, the set of levels Hij(t) for C(i, j, t) are defined, inductively, as follows: for t = 0 Hij(t)=\u2205 for t > 0 Hij(t)= 1. \u2205 if C(i, j, t)=\u2205 2. Hij(t\u22121), if \u2203`\u2208Hij(t\u22121) s.t. z\u2208[Zm, ZM ] 3. (Hij(t\u22121)\\{[Zm, ZM ]})\u222a{[Zm, z]}, if \u2203`\u2208Hij(t\u22121) s.t. 0\u2264z\u2212ZM\u2264RH and \u2200`\u2032 = [Z \u2032m, Z \u2032 M ]\u2208Hij(t\u22121), Z \u2032m\u2212z > RH 4",
            " Therefore, for each node of the graph G(A,B, t) the probability that a point belongs to a stable neighborhood is computed, and the cost map is simply defined as follows. Definition 9 (Cost Map): Let V be as defined above, let G(A,B, t) be the graph of all the neighborhoods of the fixed traversable region, and let \u0398 = (\u03a31, \u00b51), . . . , (\u03a3k, \u00b5k)) be the estimated mixture parameters with k components. The mixture g(v|\u0398) returns the probability that an observation v belongs to a stable region. Hence, the cost of the observation is 1\u2212 g(v) and of the map is 1\u2211 i ni \u2212 g(V). The algorithm is implemented for the NIFTi UGV schematically represented in Figure 2. Tests have been performed on different scenarios including stairs, slopes, ramps and rugged terrain. All tests performed were teleoperated. The operator never anticipated a path and followed the direction of the cost map, illustrated in the figures. In particular, for the experiments on the ramp (see Figures 4 and 7) the operator has used the space segmentation provided by the algorithm to trace a path, which the robot has faithfully followed. The algorithm takes as input the point cloud every 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003113_s0146411620020078-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003113_s0146411620020078-Figure2-1.png",
        "caption": "Fig. 2. Carrot chasing path following in both straight line (left) and circular (right).",
        "texts": [
            " The distance between the VTP and the UAV position projected on the path is often called virtual distance. In this research, CC is adopted to evaluate the designed controller as a final stage to guide the UAV along its desired trajectory. CC is implemented for both straight-line trajectory and the circular trajectory where both algorithms are described in the following section. The initial position of the UAV is p(xe, ye, h) with the actual heading angle \u03c8. Since the objective is to be on the path, the demanded heading angle \u03c8d guides the UAV to follow the path. Figure 2 shows a set of waypoints Wpi and Wpi+1, with the LOS angle \u03b8. The cross-track error d is a measure of the normal distance from the path. The main task of the path following algorithm is to minimize (d \u2192 0 & \u03c8, \u03b8 \u2192 0) as (time \u2192 \u221e). The designed algorithm is of three steps: first, calculating the cross-track error d, then the location of the VTP is to be updated and finally, producing the guidance command. Assume the UAV location is p with the current heading angle . The perpendicular distance from the actual position of the UAV to the path is the cross-track error d. In the loiter case, the UAV is trying to follow a circular path with origin O. The distance d is the distance between the UAV position p and O minus the circle radius r and the generated commands are to keep the UAV velocity tangent to the assumed circular path. Figure 2 is an illustration for CC assumptions to follow both the straight line and circular trajectories while Fig. 3 is a descriptive procedure f law chart to generate the required commands to follow both paths and how to switch between them. The loiter action can be either if requested at any moment or combined with the straight-line path at the designated waypoints. In the second case, the loiter sequence is initiated if the distance between the vehicle and the required-to-loiter-around waypoint is less or equal the required loiter distance, and ends when the distance to the after waypoint is minimum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003583_biorob49111.2020.9224389-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003583_biorob49111.2020.9224389-Figure1-1.png",
        "caption": "Fig. 1. Compact proton therapy system with a horizontal fixed beam-line. The patient could be treated in different positions: (a) sitting, (b) reclined, and lying position.",
        "texts": [
            " The typical size of a state-of-the-art proton therapy system ranges from 200 m2 to 600 m2 with a height of about 12 m. To increase the accessibility of proton therapy, a solution is needed to fit a proton therapy system into a conventional radiation treatment room, which is about 50 m2 with a height of 3 m. The current number of conventional X-ray radiation treatment rooms is more than 8000 while there are only 122 proton therapy rooms world wide [3]. Building a compact proton therapy system (Fig.1) which fits in a conventional treatment room would enable many more hospitals to supply proton treatment to patients. However, it is challenging to shrink a proton therapy system so dramatically from over 200 m2 (height = 12 m) to 50 m2 (height = 3 m), while maintaining the accuracy of the treatment. The positioning accuracy of the beam relative to the patient and the tumor has to be within 1 mm and 0.5\u25e6. In the conventional radiation treatment space, we need to incorporate a proton accelerator, proton beam-line and scanning nozzle, an imaging system (choices of X-ray, MRI, surface imaging, etc.) and a robot system with patient positioner and the immobilization device. The space left for the robot system is only about 4 m2. Our strategy (Fig.1) is to reduce the size and the capital cost of a proton therapy system by removing the gantry. Instead of using a 100 ton proton gantry (Fig. 2a, 2b) that bends the beam around the patient to allow for different directions of beam incidence, a horizontal fixed proton beamline is used for treatment [2], [4], [5], where the patient is moved relative to the fixed beam. A key requirement to achieve our fixed beam-line compact proton therapy is that the patient needs to be immobilized reliably and moved very precisely to different positions including the sitting (Fig. 1a), reclined (Fig. 1b), and lying positions. While others 978-1-7281-5907-2/20/$31.00 \u00a92020 IEEE 981 Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 19:14:19 UTC from IEEE Xplore. Restrictions apply. attempted this idea 20 years ago with car seat-belts, there were no soft robotics actuators at that time to control the slouching [6]. Because technology has evolved, we can now revisit this and solve that old limitation with innovative technology. In addition, horizontal beams did exist but were only used for very limited treatment sites [6], [7]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003373_10667857.2020.1810926-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003373_10667857.2020.1810926-Figure4-1.png",
        "caption": "Figure 4. Palmar and Dorsal side halves of trimmed surface in CATIA v5.",
        "texts": [
            " This conversion results in a highly accurate NURBS surface, which is then, exported to IGES format as shown in Figure 3, so that it can be further edited and modified. Creating final CAD model. Finally, the surface model is loaded in Catia v5 (Dassault Systemes, France) to obtain a CAD model of wrist immobilisation splint. The IGES surface model obtained in previous step is split into two halves, namely palmer side and dorsal side. Objective of the splint is to prevent movement of hand back and forth about the wrist. So, the palmer side is retained and dorsal side is deleted. The fingers and other unwanted portions are also trimmed as shown in Figure 4. Surface is offset by 1 mm to cater the cushion during use of the splint. A thickness of 3 mm is provided to this surface as shown in Figure 5. Two large oval shaped holes and many small holes are created on the CAD model to reduce the weight of the splint as well as to improve the comfort level of patient by promoting aeration through skin-split interface. Finally, a fillet of 1 mm and 0.5 mm radius are created to larger and smaller edges respectively and provision for Velcro strap is designed on the model as shown in Figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002979_tii.2020.2986805-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002979_tii.2020.2986805-Figure4-1.png",
        "caption": "Fig. 4. The configuration with a tilted hole and two types motion for inclined insertion. a) Tilted hole. b) In tension. c) In compression.",
        "texts": [
            " This switch condition uses force to transit from insertion, surpassing the force limitation, to withdrawal and uses a position related indicator to end the withdrawal. Combining the above controllers and the switch condition, the controller for vertical insertion with a spring is \u2206xv,t = sv (\u2206vi,t +\u2206hi,t)+(1\u2212 sv)(\u2206vw,t +\u2206hw,t) , (16) where \u2206xv,t is the manipulator\u2019s motion for vertical insertion. III. INCLINED INSERTION WITH UNKNOWN POSTURE As a necessary topic, inclined insertion occurs due to disturbance or in the case with a tilted hole as shown in Fig. 4(a), and investigation on inclined insertion can improve the insertion robustness. Commonly, inclined insertion is not desired and happens unconsciously, which requires estimating the inclination during inserting. Label the following variables: dt as the estimated inclination, xo,t as the upper object\u2019s location, and \u2206xi,t as the upper manipulator\u2019s motion of each step for inclined insertion, all at t time step. With a coil spring in the middle, the upper manipulator\u2019s movement may or may not lead to the insertion of the upper object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000592_asjc.881-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000592_asjc.881-Figure1-1.png",
        "caption": "Fig. 1. Illustration of forces and moments acting on an autonomous small-scale helicopter.",
        "texts": [
            " The position and velocity of the helicopter center of gravity are given by Pi = [x y z]T and Vi = [vx vy vz]T, respectively, with respect to the inertial frame in North-East-Down (NED) orientation (with an upper sub-index i). The helicopter angular rate vector \u03c9b = [\u03c9x \u03c9y \u03c9z]T and the Euler angle vector \u0398b = [\u03d5 \u03b8 \u03c8]T defined in the roll-pitch-yaw sequence are with respect to its body frame (with an upper sub-index b). Furthermore, \u211c( )\u0398 is the helicopter\u2019s rotation matrix from the body axes to the inertial axes and sk(\u03c9b) means the skewsymmetric matrix of the body angular rate. The principle force and moments exerted on the rigid body are illustrated in Fig. 1. \u00a9 2014 Chinese Automatic Control Society and Wiley Publishing Asia Pty Ltd The four independent inputs to this model are one lift force control, denoted by u T eT b mr b= \u2212 3 , and three directional torque controls, denoted by u L M N RM b mr mr tr T= [ ] \u2208 3. Both uT b and uM b are applied along the body frame. Qmr and Qtr represent the anti-torques on the main rotor and tail rotor, respectively. J is related to the helicopter mechanical characteristics and is expressed by J k T h k T h b mr mr b mr mr= + ( ) + ( ) \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 1 0 0 0 1 0 0 0 1 (5) where hmr denotes the height of rotor head above the helicopter center of gravity (c"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001488_amm.611.90-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001488_amm.611.90-Figure1-1.png",
        "caption": "Fig. 1 Crank Rocker Mechanism",
        "texts": [
            " The aim of this article is to create a functional model of a crank rocker mechanism in ADAMS/View software and to make its complete kinematic analysis. The aim is to investigate the movement of individual members of the mechanism and its points. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.216.129.208, University of Auckland, Auckland, New Zealand-11/06/15,05:10:39) Kinematical analysis of crank rocker mechanism We analyze the movement of point A and point B on the member 3 (Fig. 1). The member 3 is performing planar motion. The point A of the member 2 moves in a circle with radius 12 and point B of the member 4 also moves in circle with radius 14. We determine angular velocity and angular acceleration of member 2, angular velocity and angular acceleration of member 4. Crank mechanism is a simple machine used to transform linear translational (sliding) motion to rotary motion and vice versa. Although it is a simple mechanism, it has a very wide use. Kinematical analysis is shown in the four-link mechanism which is shown in Fig. 1. Kinematical analysis of the mechanism means to solve kinematical variables of the movement of the driven members with respect to kinematic variables of the movement of the driving members. The mechanism consists of a four-link crank rocker mechanism (Fig. 1). Respective lengths are: l2=120 mm, l3= 250 mm, l4=260 mm, O21O41=300 mm. Driving link O21A has a counterclockwise angular velocity \u03c921=1 (rad/s). Our task is to determine angular displacement \u03c641, angular velocity \u03c941 and angular acceleration \u03b141 of the link O41B graphically for the crank position indicated and then to create a model of a mechanism in MSC Adams/View environment. The crank O21A rotates around point O21, motion of the connecting member AB is a general planar motion and member O41B rotates around point O41",
            " Simulation of the crank rocker mechanism using MSC Adams/View The given crank rocker mechanism was modeled in MSC ADAMS/View and the initial parameters were provided [2],[3]. In the initial window of the program Adams we set data folder, name of the project, units and the working grid. We created individual bodies of the mechanism. We selected the rigid body link from the Toolbox. We defined the geometry, length, width and depth. The lengths of the four links crank rocker mechanism are: l2=120 mm, l3= 250 mm, l4=260 mm, O21O41=300 mm. Driving link O21A (Fig. 1) has a counterclockwise angular velocity \u03c921=1 rad/s. Member 2 [2], [3] must be created and link tool in Main Toolbox must be selected. Move the cursor to (0, 0, 0) position and member 2 is created a named automatically. Then, move the cursor to (300, 0, 0) position and create the member 4. The member 4 should be rotated about the point from horizontal position to counterclockwise position, rotated by 114\u00ba (Fig. 5a). Then, connect the member 2 and 4 with member 3 (Fig. 5b). Next, joints were created between the members and the ground (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001469_s1023193514120027-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001469_s1023193514120027-Figure7-1.png",
        "caption": "Fig. 7. Interferences study for Acarbose to measure the AG enzyme inhibition.",
        "texts": [
            " Therefore, Tebengau plant extracts may be used as a good medic inal based candidate for the development of antidia betic drugs. In order to study the selectivity of the method, 1 mg/mL of Acarbose solution were added in the con trol solution and recorded the voltammograms of CV. Then, the common interferences such as methanol, ethanol, Na+, K+, Ca2+, Mg2+ were mixed with Acar bose and control solution and recorded voltammo grams. There was no other peak response appeared and same inhibition was observed as without interfer ences which is shown in Fig. 7. Thus indicating the excellent selectivity of the MWCNTs paste electrode to measure the antidiabetic potential of medicinal plants as well as commercial antidiabetic drugs. To study the repeatability of the fabricated MWCNTs paste electrode, 1.0 mM PNPG solution and the fabricated electrode were used every two hours within a day. The electrode showed an almost stable peak current during every measurement obtained using the cyclic voltammetric method at constant temperature. After 15 days, the voltammograms of PNPG were measured under the same conditions with the same electrode and recorded"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002290_b15961-60-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002290_b15961-60-Figure4-1.png",
        "caption": "Figure 4. Demonstrator.",
        "texts": [
            " But in sheet metal forming, tool load in terms of compression and tension is considerably higher and more demanding, setting a new challenge for laser beam melted tooling. The aim of the presented Innovation Alliance project was the development and manufacturing of tool inserts with an optimized cooling system to improve the resource efficiency in hot sheet metal forming. Therefore, thermo-fluidic simulation and laser beam melting were used. After investigating current mass production, the project partners have jointly developed a representative demonstrator (see Figure 4). To enable easy transfer of the project\u2019s results into mass production, the demonstrator\u2019s geometry is very similar to a serial component. The design reflects a typical hot forming component and its difficulties and potential problems. It incorporates geometric features such as curved surfaces and cavities to demonstrate limitations of conventional, deep hole drilled cooling channels in terms of rapid and homogeneous cooling of the sheet metal component. The tool design and the cooling system design were done based on conventional manufacturing methods such as milling and deep drilling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003016_j.sna.2020.111883-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003016_j.sna.2020.111883-Figure2-1.png",
        "caption": "Fig. 2. The structure of linear motor.",
        "texts": [
            " The first part is the linear oscillating otor, the second part is the cylinder which has two chambers harging or discharging oil. The third part is a spool valve working s a rectification valve. Movers of three parts are linked together nd driven by the linear oscillating motor for linear reciprocating esonant motion. Because the inclination torque is ignorable, the riction is much lower compared with its counterparts driven by otary motors, and high reliability can thus be achieved. As shown in Fig. 2[17], one linear oscillating motor is designed or the linear pump as its resonance status can improve the workng efficiency greatly. Silicon steel is laminated circumferentially o eliminate eddy current in the stator. Halbach magnet array is ounted on the mover iron to generate high flux density. Two prings are mounted on both sides of the stator for motions at esonance. The cylinder and the spool valve are integrated into one ydraulic block. The cylinder has two chambers separated by the iston, and the two chambers charge and discharge alternatively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure2.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure2.1-1.png",
        "caption": "Fig. 2.1 Schematic drawing of linear machines with dual Halbach array",
        "texts": [
            " Based on PM arrangement, magnetic field distribution in the machine is formulated with Laplace\u2019s and Poisson\u2019s equations analytically. Numerical result from finite element method is utilized to analyze and observe flux variation in three-dimensional space of the machine. The numerical result validates the analytical models. The obtained analytical model could be used for analysis of output performance and control implementation of electromagnetic linear machines with similar structures. The schematic structure of the proposed PM tubular linear machine is illustrated in Fig. 2.1. The mover consists of winding phases mounted in the holds, whereas the stator is composed of dual Halbach magnet arrays that enclose the windings on internal and external sides. 2.2 System Structure and Operating Principle 19 The polarization pattern of each layer of the PMs is arranged in Halbach array. As indicated in Fig. 2.2, the Halbach array consists of alternatively magnetized PMs in radial directions separated with horizontally magnetized PMs. This array offers one impressive benefit, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002003_ajplung.00117.2013-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002003_ajplung.00117.2013-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the experimental procedure. A trachea was excised by dissection, and a specimen of about 3 3 mm was placed between the slide glass and the coverslip.",
        "texts": [
            " To understand the origin of flow fluctuation, we first measured the distribution of ciliated cells in the trachea and their ciliary motion. Next, we measured the in-plane flow field at different heights of observation using a confocal micro-PTV system. This system enabled us to accurately measure the flow with high resolution in time and space. Finally, we quantified the effect of flow fluctuation on bulk flow by evaluating the Peclet number (cf. Eqs. 7 and 2) of the system. Preparation of samples. Experiments were conducted with approval by the Animal Ethics Review Board of Tohoku University. Figure 1 illustrates the procedure for sample preparation. Tracheas were obtained exclusively from wild-type mice (Crlj:CD1), 4\u201316 wk old. The mice were killed by cervical dislocation. The trachea was then dissected from the larynx to the bronchial main branches and doused with L-15 medium (GIBCO) 10% fetal bovine serum (FBS; Thermo Scientific). Muscle and vascular tissues were removed from the trachea in cold medium. The trachea was opened longitudinally, and 3-mm square specimens were excised. The exposed tracheal lumens were soaked in L-15 medium including 5 mM dithiothreitol (DTT; Fluka) to gently remove the original mucus and contaminating Address for reprint requests and other correspondence: T"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000477_imece2014-37638-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000477_imece2014-37638-Figure1-1.png",
        "caption": "FIGURE 1: Coordinate system used with the models",
        "texts": [
            " A least-squares calibration achieve an error below 1.37%. The total error in the measurement of drag, side force, and yaw moment coefficient was calculated to be \u00b11.21%. This value was calculated using the errors due to individual factors found in Table 2. The yaw angle uncertainty was determined to be \u00b1 0.09\u25e6 from the micro-stepping capability of the stepper motor. Error due to nonlinear interactions of the model and the supports was verified to be negligible using the mirror mount procedure [16, 1] Figure 1 shows the coordinate system used with the models. Please note that the sense of yawing moment as presented in the figure (and used in the data) is opposite to the usual right-hand rule for moment. This difference arises from the use of the load cell and reaction system developed for these experiments. Results from 4 models are discussed in this paper. The first is a cylinder model of aspect ratio (L/D) 1, made from a cardboard tube. Results were shown in Ref. [1] and are not repeated here. The second is a cuboid with scaled dimensions from the CONEX container geometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002813_ilt-06-2019-0239-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002813_ilt-06-2019-0239-Figure5-1.png",
        "caption": "Figure 5 Three different shapes of dimples (a) curvature parallel to the sliding direction, (b) curvature perpendicular to sliding direction",
        "texts": [
            " Saleh et al. Industrial Lubrication and Tribology Volume 72 \u00b7 Number 3 \u00b7 2020 \u00b7 405\u2013414 pressure and the load-carrying capacity. Evidently, this difference is quite obvious at a high-pressure regime, as there is no effect at low or zero pressure zone. One of our interests in this work is to study the impact of texture dimple\u2019s height curvature on the journal bearing under THD conditions. Three shapes of surface curvature (rectangle, triangle, and ellipse) have been used in this investigation as shown in Figure 5. It has been noted that the dimples with rectangle height have the overall best performance. The surface with rectangle height dimples (D5, D-7, D-8) or rectangular grooves (G-5, G-7, G-8) has load-carrying capacity higher than that of the other two shapes. The friction force produced from ellipse dimples is higher than that of the triangle dimples but still less than that of the rectangle dimples. On the other hand, the lubricant temperature increases in the case of rectangular dimples and the altitude angle decreases as illustrated in Figure 6. According to the results of all cases, it has been proved that the rectangular shape for texture in channel or dimples form gives the best THD performance compared with the ellipse or triangle channel. The effect of the relation between the direction of curvature and the sliding direction has been studied. As demonstrates in Figure 5 there are two cases. In the first case, the curvature direction is parallel to the sliding direction as (D-1, D-3, D-7, D-9, D-11), while in the second, it is perpendicular to the sliding direction as (D-2, D-4, D-8, D-10, D-12). The textured surface with height curvature perpendicular to the sliding direction have a value of pressure, load carrying capacity and friction force higher than the other curvature direction. Otherwise, the temperature of the lubricant in the case of perpendicular curvature is higher in the region of high pressure while the temperature in the case of parallel curvature is higher in the region of low or absent pressure as shown in Figure 7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure2-1.png",
        "caption": "Fig. 2 Schematic of hair and socket structure and the receptor neuron arrangement",
        "texts": [
            " Two long antennalike appendages, called cerci, are the sensory organs for the cercal sensory system. Each cercus is approximately 1 cm long in an adult cricket and is covered with 500\u2013750 mechanosensory filiform hairs. These hairs range from 50 lm to 2 mm in length [4]. Figure 1 shows the location of the cerci of a cricket. The base of each hair is attached to the cercus and constrained within a cuticular socket, which also contains the active region of a spike generating sensory neuron [7], shown schematically in Fig. 2. In the resting state with no air-current stimuli, the hair projects up through the center of the cuticular socket, and its associated receptor neuron generates action potentials at a baseline rate of approximately 30 Hz [4]. Any air movement causes a deflection of the receptor hair away from its rest position. Each filiform hair is constrained at its base to move through a single plane by a hingelike support. Movement in one direction through a hair\u2019s movement plane excites the associated receptor neuron to generate action potentials above its baseline activity rate, and movement in the other direction 1Corresponding author"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000518_eeeic.2015.7165426-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000518_eeeic.2015.7165426-Figure2-1.png",
        "caption": "Fig. 2. Generator 2D model",
        "texts": [
            " It is a salient poles three-phase synchronous generator with DC excitation winding. The main data are: outer diameter of the stator core, 106.5 mm, inner diameter, 70 mm , length, 75 mm, outer diameter of the rotor, 67.3 mm, pole pairs number, 2, stator slots number, 36, stator tooth width, 2.75 mm, turns number 133, phase voltage, 230 V, phase current, 0.7 A, rated power 0.37 kW. The stator winding is in a single layer having three slots per pole and phase. The used software, MagNet, is dedicated to the numerical calculation of low frequency magnetic field. The 2D model, Fig. 2, is used to perform transient with motion analysis. The stator phases are presented in different colors. Because the model length is only half of the generator length, the windings parameters and voltages will be reduced accordingly. The magnetic core is considered nonlinear, the conductivity being set to zero (iron losses being taken into account by Steinmetz equation based on frequency and magnetic flux density). External circuit elements are connected to the numerical model as follows: a resistance for each phase of generator, its value being chosen so that the model has the same resistance as the real machine; an inductance for each stator phase representing the end winding inductance; a resistance and an inductance for field winding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000774_cjme.2013.03.573-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000774_cjme.2013.03.573-Figure3-1.png",
        "caption": "Fig. 3. Coordinate systems for RTCA",
        "texts": [
            " The origin of the coordinate system is at the apex of the CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7575\u00b7 pitch cone. X-axis is aligned with the gear axis of rotation, and directed to the big end of the gear. Z-axis is along radial direction of the gear and is directed away from the to-be-measured real tooth surface. Y-axis is determined according to the right hand rule. The center point P of the grid lies in O-XZ plane. Five coordinate systems are applied to describe the relative position of the meshing tooth surfaces as shown in Fig. 3. Coordinate system Sm is rigidly connected to the machine frame. Movable coordinate systems S1 and S2 are rigidly connected to the pinion and gear, respectively. Auxiliary coordinate systems S10 and S20 are used to describe the rotation of the pinion and the gear, respectively. \u03b3 is the shaft angle. 1\u03c6 and 2\u03c6 are the rotation angle of the pinion and gear, respectively. RTCA is based on the fact that the real tooth surfaces of the pinion and the gear are considered as being in continuous tangency which implies mathematically as m1 1 1 1 1 m2 2 2 2 2 m1 1 1 1 1 m2 2 2 2 2 ( , , , ) ( , , , ) 0, ( , , , ) ( , , , ) 0, ( , , ) 0, 1, 2, i i i i X Y Z X Y Z X Y Z X Y Z f X Y Z i \u03c6 \u03c6 \u03c6 \u03c6         r r n n (1) where the lower index 1 represents the pinion, and the lower index 2 represents the gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001040_j.mporth.2016.05.014-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001040_j.mporth.2016.05.014-Figure9-1.png",
        "caption": "Figure 9 (a) Work done\u00bcFd. (b) Work done\u00bcF cos qd. (c) Work done in lifting weight against gravity\u00bcmgh.",
        "texts": [
            " Then, by moving the limbs closer to the axis of rotation, the moment of inertia of the skater is reduced. If the frictional effects of the ice are ignored, then no external moments or forces (other than gravity) are acting on the skater and therefore their angular momentum (\u00bcIu) will remain constant. Thus if I decreases, the angular velocity u must increase and the skater will rotate faster. Work is defined as the product of a force and the distance moved by the point of application of the force in the direction of the force, shown in Fig. 9. Energy is defined as the capacity to do work. Work and energy are measured in units of joules (J) where 1 J\u00bc1Nm. Potential energy and kinetic energy are the most common forms of energy encountered in biomechanics. Potential energy relates to the position of an object relative to an initial point and equals the work done in moving it from the initial point to its present position, for example, lifting a weight against gravity. By lifting a mass m through a height h, the work done against the force of gravity is equal to the force acting on the object (\u00bcmg) multiplied by the distance moved along the line of action of the force: Kinetic energy relates to the velocity of an object and is the work done in increasing the velocity of the object from zero to its present value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002109_j.proeng.2015.12.162-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002109_j.proeng.2015.12.162-Figure1-1.png",
        "caption": "Fig. 1. The MGTP rotor",
        "texts": [
            " Peer-review under responsibility of the organizing committee of the International Conference on Industrial Engineering (ICIE-2015) Nomenclature MGTP Micro gas turbine plant TCR Turbocharger rotor SGR Rotor of the starter-generator SGR Rotor of the starter-generator FEM Finite element method Low-powered micro gas turbine plants are utilized in industrial enterprises, medical centres; on main gas lines, oil pipe lines, gas distribution stations; in power-hungry regions of Extreme North, Siberia, Far East; to replenish electrical shortage caused with natural disaster s and other emergency situations. The most important part of the MGTP is rotor; its operational frequency is 65,000 rpm. It consists of turbocharger rotor (TCR) and starter-generator rotor (SGR) connected by elastic coupling (Fig. 1). One of the requirements to high-speed rotor is exclusion of falling its critical speeds in the range of \u00b130% of operating speed (45,500\u201384,500 rpm) [5]. Thus, there rises an objective to develop a rotor with critical speeds, which do not fall in the prohibited range. Works [1, 2, 6, 7, 17-20] are devoted to the same matter. The evaluation is performed through the finite element method (FEM) in the package Ansys Workbench. Natural frequencies were calculated at free-free boundary conditions. Natural frequencies of bending vibrations of the rotor in the range from 0 to 120,000 rpm (2,000 Hz) are represented in table 1, frequencies and shapes of torsional and longitudinal vibrations are not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000250_iros40897.2019.8968074-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000250_iros40897.2019.8968074-Figure6-1.png",
        "caption": "Fig. 6. Measurement setup for bending modulus. The setup mainly consists of an electronic scale, an X stage with a micrometer, and a rod. The reaction force when bending the prototype was measured by the electronic scale with a rod on. The prototype is set to the X stage with a micrometer and pushed down to the rod by the micrometer.",
        "texts": [
            " After one hour, the process is repeated; however, this time the liquid is flowed in the opposite direction. After two iterations of this two-stage process, the entire process is complete. This subsection describes a measurement of bending modulus as well as an electrolysis experiment as a basic evaluation of the soft PEFC tube. 1) Measurement of bending modulus: Young\u2019s modulus is measured in this subsection to confirm that the developed PEFC tube is soft enough for the practical realization of the proposed concept. The measurement setup is depicted in Fig. 6. The setup mainly consists of an electronic scale (AJ-12K, SHINKO DENSHI CO.,LTD.), an X stage with a micrometer, and a rod. The reaction force when bending the prototype was measured by the electronic scale with a rod on. Note that the length of the test piece is 14.3 mm. The prototype is set to the X stage with a micrometer and pushed down to the rod by the micrometer. The results of the measurement are shown in Fig. 7, which plots the relationship between displacement and bending force. From the results in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002233_1350650116661071-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002233_1350650116661071-Figure5-1.png",
        "caption": "Figure 5. Upper and lower specimens.",
        "texts": [
            " Measuring wear coefficient of cam material in boundary lubrication According to the cam\u2013follower pair of the valve train in a diesel engine, the wear coefficient of the cam material in boundary lubrication is measured using the transient friction and wear testing machine, as shown in Figure 4. The upper test specimen of the cam material is fixed on the main axel, which rotates at a given speed and applies a certain force on the lower test specimen of the tappet material, which is fixed on the lower set. The upper and lower specimens used in this test are shown in Figure 5. Their materials, hardness, and RMS roughness of the specimens are listed in Table 1, which are the same as those of the cam and the tappet in the valve train of the engine studied in this paper. During the test, the boss of the upper specimen maintains contact with the lower specimen, thus the wear coefficient of the cam material is obtained by measuring the height reduction of the boss. Under boundary contact conditions, the critical value of the oil film thickness, hb, is given by10 hb \u00bc 0:3 0 \u00f05\u00de and 0 is composite surface roughness parameter, which is calculated using the RMS roughness of the upper specimen and lower specimen surfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure2-1.png",
        "caption": "Fig. 2. An equivalent articulated manipulator to the wheeled mobile manipulator in Fig. 1.",
        "texts": [
            ", the one for mobile platform and the other for its on-board arm), the proposed approach describes the motion of a WMM only by an equivalent single articulated manipulator. We shall consider two representative types of mobile platforms, from a simple differential-drive type to a more complicated car-like type in order to apply the proposed modeling approach - however, in principle, the applicability of the approach is not limited to these. 3.1. Differential-drive mobile platforms WMMs with differential-drive mobile platforms as shown in Fig. 1 can be physically transformed into ordinary articulated manipulator systems as shown in Fig. 2, where an imaginary revolute joint is anchored at the center of the wheel axle, Op, and an imaginary prismatic joint is attached atop the revolute joint. The structure of the transformed system exhibits the velocity kinematics identical to that of the original mobile manipulator. A systematic way to carry out the transformation is summarized as follows: Step 1: Fix a virtual revolute joint to ground at the current center position of the axle of the driving wheels, Op. Step 2: Attach a virtual prismatic joint with length D atop the virtual revolute joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002237_978-3-319-29357-8_59-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002237_978-3-319-29357-8_59-Figure1-1.png",
        "caption": "Fig. 1 The grasp examples used for learning: a pinch grasp, b spherical grasp, c power grasp, d hook grasp",
        "texts": [
            " The whole procedure of learning and generating grasps for novel objects consists of four stages [1]: \u2219 grasp training on a small set of example grasps: the Contact Model (Mi) for each link and the Hand Configuration Model (C) for whole gripper are generated, \u2219 estimation of PDF of gripper links\u2019 poses for a novel object (the Query Density Qi), \u2219 sampling the Query Density Qi and the Hand Configuration Model C: a set of grasps is generated, \u2219 grasps optimization and selection The learning stage requires a small number of example grasps (Fig. 1) and even one example of each grasp type is sufficient. In this section we introduce some of the concepts from [1] and we propose a number of modifications. The following symbols are used for clarity: 1.   p\u2014translation vector of the origin of a frame  expressed in the reference frame , 2.   q\u2014orientation of a frame  with respect to (w.r.t.) the reference frame . 3. P\u2014position of the point P expressed in the reference frame , Object Model is a PDF of object\u2019s surface. It is used to estimate the shape and surface features of an object",
            " The set of test objects included object models from the original work [1] (a kettle, a mug, containers) and some shape primitives such as a cuboid and cylinders. The gripper used for tests was BarrettHand BH-280. It is a 4-DOF under actuated gripper with 8 joints coupled in pairs. Each of the three fingers has two flexion joints: the proximal joint and the coupled distal joint. Two fingers together have one additional DOF, i.e. the two coupled spread joints. The grasp examples for the BarrettHand gripper used in the learning step are shown in Fig. 1. The Contact Models together with the Configuration Model were generated for each grasp example. They were used to generate the Query Densities for the Object Model of a novel object. Then, the Query Density was used to generate a set of grasps that were later optimized as described in [1]. The highest-ranked grasps were selected for the collision checking procedure. If the gripper in the initial configuration for the given grasp was in collision with the environment, the pose and configuration were randomly altered in some small range several times"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002856_s00422-020-00821-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002856_s00422-020-00821-1-Figure1-1.png",
        "caption": "Fig. 1 The proposed biomechanical model. a The whole biomechanical model. bThe local magnification of the left forelimb. The shoulder joint (SHj) and hip joint (Hj) are regarded as Hooke joints. The body joint (Bj), an elbow joint (Ej), and knee joint (Kj) are regarded as revolute joints. For the limbs, the filledmuscles represent the flexormuscles, and the blank muscles represent the extensor muscles. In the body, the filled muscles represent the left muscles, and the blank muscles represent the right muscles. \u201cMuscle of LR\u201d indicates the muscles that control the leg-raising, and \u201cMuscle of LS\u201d indicates the muscles that control the forward and backward swing of the limbs",
        "texts": [
            " Section 3 constructs a spinal locomotor network model of the salamander based on the improved biomechanical model, and designs the control methods that can maintain the biomechanical model\u2019s posture, control the initial swing order of its forelimbs, generate different walking turning and turning on the spot, and implement switches between gaits. Section 4 evaluates the properties of gait generation and transition of the new spinal locomotor network model through simulation. Section 5 discusses the proposed spinal locomotor network model of the salamander, the gait control methods of the spinal locomotor network model, and the obtained simulation results. Lastly, Sect. 6 concludes the paper and gives an outlook for our future work. The improved biomechanical model proposed in this paper is shown in Fig. 1. Except for the shanks of the forelimbs, the proposed biomechanicalmodel is the same as the biomechanicalmodel introduced in Liu et al. (2018). The proposed model body consists of ten rigid links, as shown in Fig. 1a, and it is 25 cm long. The body links are connected by a single-degreeof-freedom revolute joint with the axis perpendicular to the ground at the initial posture of the biomechanical model. The shoulder joint, which is located between the forelimb thigh and body links, is regarded as a Hooke joint with the axes perpendicular to the ground and parallel to the body axis at the initial posture of the biomechanical model. The hip joint, which is located between the hindlimb thigh and body links, is similar to the shoulder joint. The knee joint, which is located between the thigh and shank links of the hindlimb, is regarded as a revolute joint with the axis parallel to the body axis at the initial posture of the biomechanical model. The biomechanical model of the forelimb shanks shown in Fig. 1b is built based on the characteristics of the salamander forelimbs. The elbow joint, which is located between the thigh and shank links of the forelimb, is regarded as a revolute joint with the axis perpendicular to the plane passing through the shank and thigh links of the forelimb. Based on the observation, such a model structure is advantageous in modeling the terrestrial turning of the salamander. Following the improved biomechanical model, we use ADAMS 2013 software to build the biomechanical model for dynamic simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure5-1.png",
        "caption": "Fig. 5. Detail view of the fabricated pump module",
        "texts": [
            " The battery module contains lithium polymer battery-cells combined into packs, safety electronics and voltage converters to allow monitoring of charge and discharge. An image of the layout of the module is shown in Fig. 4. The pump module was designed to provide the pneumatic actuator of the support system for the necessary centration. Also, it maintains 0.6 MPa of differential pressure to operate external pressure (5.0 MPa). DAS (Data Acquisition System) of NDE module, sub controller and hub PCB for communicating with NDE system and tractor robot were integrated (Fig. 5). The MFL system was designed to minimize friction between magnetizer and pipe wall, and maximize defect detectability. Also, the performance of the MFL system was verified, and it satisfies the specification of POF (Pipeline Operate Forum) standard. Figure 6 shows the fabricated MFL system for inspection of unpiggable pipelines with the PIBOT. The MFL system includes link system and collapsing system to negotiate obstacle structures. Also, the MFL system has DAS system to acquire MFL signals. Using the pull-rig test facility, MFL signals, which were line scanned from the gas pipeline with artificial defects, were acquisitioned as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001117_s1064230713050109-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001117_s1064230713050109-Figure1-1.png",
        "caption": "Fig. 1. The kinematic scheme of the wheeled robot.",
        "texts": [
            "1134/S1064230713050109 820 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 52 No. 5 2013 PESTEREV applying the method proposed in [12]. Advantages and drawbacks of the three feedbacks are discussed. Comparison of the control law derived in this paper with two above mentioned feedbacks demonstrates its unquestionable advantages. 1. PROBLEM STATEMENT The wheeled robot considered in this work is a vehicle moving without lateral slippage with two rear driving wheels and front wheels responsible for steering the platform. The kinematic scheme of the robot is depicted in Fig. 1. In the planar case, the robot position is described by two coordinates of the point located in the middle of the rear axle (further referred to as target point) and the angle describing orientation of the platform with respect to a fixed reference system x0y. Every point Xp of the platform has its own instantaneous velocity vector . Vectors orthogonal to the instant velocities intersect at the instan taneous center of velocities X0. The instantaneous angular rate of the platform rotation satisfies the rela tionship Hereinafter, denotes the Euclidean vector norm. The condition that the wheels move without lateral slippage means that the vectors of instantaneous velocities of the axles' endpoints are collinear to the planes of the wheels and that the normals to these vectors intersect at the point . For the rear axle, the instantaneous center of velocity coincides with the instantaneous center of cur vature. Particularly, for the target point, the value is the instantaneous radius of curvature of the trajectory (the dashed line in Fig. 1) circumscribed by the target point . The reciprocal to the radius is the instantaneous curvature . It can be seen from Fig. 1 that the turning angles of the two front wheels are different and are related to the curvature of the trajectory described by the target point by the equations where is the distance between the left and right front wheels and is the wheelbase. The trajectory cur vature and the steering angles are considered positive when the platform turns left (counterclockwise), as shown in Fig. 1. The relationships between the steering angles and the trajectory curvature at the target ,( )c cx y cX \u03b8 pv \u03b8 \u03b8 = \u2212 . | | / 0p pX Xv \u22c5 0X \u2212 0cX X cX \u2212/ 01 cX X \u03ba tan tan / / 1 2 1 2 1 2 L L H H \u03ba \u03ba \u03b4 = , \u03b4 = , \u2212 \u03ba + \u03ba H L JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 52 No. 5 2013 A LINEARIZING FEEDBACK FOR STABILIZING A CAR LIKE ROBOT 821 point allows us to simplify the model and, instead of two angles and , introduce an \u201caverage\u201d angle by the equation The angle may be interpreted as the turning angle of a single imaginary front wheel located on the central line of the platform at the distance from the rear wheels of a three wheeled platform that is kinemati cally equivalent to the considered four wheeled vehicle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001737_978-3-319-23204-1_11-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001737_978-3-319-23204-1_11-Figure1-1.png",
        "caption": "Fig. 1 A badminton racquet integrated with an acceleration signal acquisition module.",
        "texts": [
            " In the remaining of this study, Section 2 describes the movements in badminton and their signal processing procedures. It also describes the algorithms of signal acquisition, the features\u2019 extraction, and classification. Section 3 shows the experimental results and concludes our work in Section 4. In assessing the effectiveness of badminton swing,we developed a badminton racquet signal acquisition system, which includes an inertial acceleration signal acquisition module equipped with a three-axis accelerometer, as shown in Fig. 1. The racquet is parallel with the Z axis of the triaxial accelerometer, and perpendicular with the Y-axis, when the players perform swing movement. The acceleration signals caused by changes will be recorded in the memory card in the module. There are three common badminton actions, which are clear, drop, and smash. In fact, these three actions have very close decomposed steps. Therefore, the clear action is considered as the most fundamental action in racquet swing. Generally, good players in dealing with the same kind of shuttle pitch by the strokes will be very similar since this is already accustomed to the action for him"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003050_s00033-020-01309-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003050_s00033-020-01309-5-Figure1-1.png",
        "caption": "Fig. 1. Configurations of a radially symmetric accreting hollow cylinder on a rigid substrate (inner disk in gray). Top two disks: material and deformed configurations at time t1, 0 < t1 < tf . Bottom left two disks: material and deformed configurations at the final time tf . Bottom right disk: when accretion is completed one unloads and lets the cylinder cool down. The result is a residually stressed configuration",
        "texts": [
            " In our numerical examples, we consider the simple model K = KG , where K is the heat conduction coefficient, while we call D = K cE\u03c1 the diffusivity coefficient. Therefore, the heat equation (17) is simplified to read5 Div(KG dT) = \u03c1cE T\u0307 \u2212 1 2 T \u2202S \u2202T :C\u0307 \u2212 \u03c1R. (18) 4In Appendix D we discuss the derivation of this inequality. 5Note that G dT = Grad T . In this section, we present a simplified formulation of the thermoelastic problem for the radially symmetric accretion of an infinitely long hollow cylinder (see Fig. 1). We assume that the material is initially added on the outer surface of a rigid and infinitely long cylindrical substrate of radius R0 and that the growing cylinder is sitting in an ambient temperature Ta. We assume that stress-free cylindrical layers of new material are continuously and uniformly formed on the outer boundary of the cylinder at a temperature Tm\u2014its melting temperature\u2014greater than Ta. The growth velocity is assumed to be normal to the growth surface and has magnitude u(t). The added rings are made of the same homogeneous isotropic incompressible material, with a uniform isotropic coefficient of thermal expansion \u03b1(T ), and a uniform isotropic coefficient of heat conduction K(T )",
            " One can hence define its inverse map \u03c4 = S\u22121 assigning to each ring at R \u2265 R0 its time of attachment \u03c4(R) as defined in Sect. 2.1. Assuming that accretion starts at t = 0, it follows that \u03c4(R0) = 0 and equivalently that S(0) = R0. We define the material growth velocity U(t) = S\u0307(t), which yields S(t) = R0 + \u222b t 0 U(\u03bd)d\u03bd . In what follows, we choose S(t) such that U(t) = u(t),6 which means that in the time interval [t, t + dt] the external radii of both the material and spatial disks grow by the amount u(t)dt (see Fig. 1). Hence, S(t) = R0 + \u222b t 0 u(\u03bd)d\u03bd. Kinematics of the accreting cylinder. We assume that the growing cylinder deforms in a radially symmetric fashion, i.e., we take motions \u03d5t of the kind (r(R, t),\u0398). The deformation gradient reads F (R, t) = [ r,R(R, t) 0 0 1 ] , (20) where a subscript comma denotes partial differentiation, e.g., r,R(R, t) = \u2202r \u2202R (R, t). As was mentioned earlier, as the growth surface R = S(t) is traction-free, Q = F\u0304 , and hence, u = F\u0304U . Therefore, the material growth speed U and the spatial growth speed u are related as u(t) = r,R(S(t), t) U(t). (21) This is equivalent to assuming the absence of traction on the outer boundary, see [49]. Having already assumed U(t) = u(t), Eq. (21) gives us r,R(S(t), t) = 1, or r,R(R, \u03c4(R)) = 1, (22) so that the frozen deformation gradient reads F\u0304 = I; see Sect. 2.1. We introduce the notation s(t) = r(S(t), t); see Fig. 1, and r\u0304(R) = r(R, \u03c4(R)), so that one has s = r\u0304 \u25e6 S, or r\u0304 = s \u25e6 \u03c4 . Finally, the rate of change of the spatial radius of the growing cylinder s(t) = r(S(t), t), representing the radial and the only nonzero component of the total velocity w, is written as s\u0307(t) = d dt [r(S(t), t)] = r,R(S(t), t)S\u0307(t) + r,t(S(t), t) = S\u0307(t) + r,t(S(t), t), (23) where r,t = \u2202r/\u2202t is the radial and only nonzero component of the standard velocity v. From (23), it follows that the velocity of the accretion boundary is the result of two contributions: S\u0307(t) = U(t) = u(t), due merely to accretion, and the standard velocity r,t(S(t), t) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure13-1.png",
        "caption": "Fig. 13. The Singular configuration I.",
        "texts": [
            " In both plane-symmetric Mode I and plane-symmetric Mode II, there are some singular configurations, such as recon- struction configuration and parallel configuration. Starting from these singular configurations, the mechanism can reach configurations which are different from the configurations of plane-symmetric Mode I and plane-symmetric Mode II. When the joint axes of R joints J R2 , J R4 , J R6 , and J R8 are parallel, and the joint axes of R joints J R1 , J R3 , J R5 , and J R7 intersect at a point ( Fig. 13 ), the configuration is a bifurcation of plane-symmetric Mode I and plane-symmetric Mode II. Since the axes of the R joints J R2 , J R4 , J R6 , J R8 are parallel, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration I ( Fig. 13 ), when the joint axes of R joints J R2 , J R4 , J R6 , and J R8 are always parallel and in the same directions during the movement, the mechanism makes a planar motion, which is called plane-motion Mode III. Its typical configurations are shown in Fig. 14 . When the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R1 and J R5 are parallel, and the joint axes of the R joints J R3 , J R7 are also parallel, as shown in Fig. 15 . This configuration is a special configuration of Mode I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001215_icmech.2015.7084001-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001215_icmech.2015.7084001-Figure1-1.png",
        "caption": "Fig. 1. Side view of electric motorcycle",
        "texts": [
            " Vehicle body part of dynamics From Fig. 2, the vehicle body can be regarded as the inverted pendulum. So, the dynamics of camber angle (13) can be derived by solving the Lagrange equation [5]. A\u03c6\u0308 = B sin\u03c6\u2212 C\u03b8\u0307 \u2212D sin \u03b8 + T dis \u03c6 (13) The variables are defined as (14)-(17) A = IX +Mh2 (14) B = Mgh (15) C = MhV L1 L cos \u03b8 cos\u03c6+ Jtrf\u03c9f (16) D = Mh V 2 L2 (L1 cos \u03b8 + L2) cos\u03c6 + (Jtrf\u03c9f + Jtrr\u03c9r) V L +Mh L1 L cos\u03c6V\u0307 (17) The relation between \u03b8 and \u03c8 are given by (18). \u03c8\u0307 = V L sin \u03b8 (18) The parameter of the figure 1 is shown as the Table I. In the (6) and (7), \u03bci(\u03bbi) is the friction coefficient, which is the ratio between the driving force and the normal tire force, depends on the road condition and wheel slip. The proposed slip ratio considering camber angle during driving \u03bbdrv i and during braking \u03bbbrk i are shown as (19) and (20) by using (18) [6]. \u03bbdrv i = r\u03c9i \u2212 ( V + r\u03c8\u0307 sin\u03c6 ) r\u03c9i = Lr\u03c9i \u2212 V (L+ r sin \u03b8 sin\u03c6) Lr\u03c9i (19) \u03bbbrk i = r\u03c9i \u2212 ( V + r\u03c8\u0307 sin\u03c6 ) V + r\u03c8\u0307 sin\u03c6 = Lr\u03c9i \u2212 V (L+ r sin \u03b8 sin\u03c6) V (L+ r sin \u03b8 sin\u03c6) (20) In this section, the details of the control methods are described"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003208_s43236-020-00116-5-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003208_s43236-020-00116-5-Figure3-1.png",
        "caption": "Fig. 3 Forces acting on a vehicle",
        "texts": [
            " Therefore, optimal reference value of the stator flux is obtained: Figure\u00a02 shows the proposed DTC block scheme for EVs. The difference between the proposed and conventional DTC methods is the use of the optimal stator flux block obtained by Eq.\u00a0(24). Although the reference stator flux is constant in the conventional DTC, it varies in the proposed optimal DTC to minimize the losses of the IM. (23)K2 = RsRFeL 2 r + sL 2 m ( L2 m + L2 r ) L2 m RFeL 2 r . (24) opt s = Ls Lm \u221a\u221a\u221a\u221a \u221a( opt dr )2 + ( 2 3 Lr P )2 ( T ref e opt dr )2 . 1 3 A vehicle under the influence of various forces is shown in Fig.\u00a03. The propelling force for the vehicle is determined by Eq.\u00a0(25) [13] where Fte is the traction force, Frr is the rolling resistance force, Fhc is the hill climbing force, Fla is the acceleration force, and Fad is the aerodynamic drag force. A Li-ion battery is preferred in this EV application, since it provides high power and energy densities. During the discharge of the battery, the voltage is calculated by (26) where Ebat is the battery voltage, E0 is the battery constant voltage, K is the polarization constant, i* is the filtered current, i is the battery current, it is the actual battery charge, Q is the maximum battery capacity, A is the amplitude of the exponential area voltage, and B is the exponential area time constant inverse"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure10-1.png",
        "caption": "Fig. 10. (a) Input\u2013Output direct contact mechanism with two tori in point contact, (b) substituted RSSR mechanism with spherical joint centers K and Ko chosen at (200,0,100) and (0,200,0) respectively.",
        "texts": [
            " This means that the centers of the substitute-connection cannot be obtained using only the curvature values obtained by mere examination of the profiles of the contacting surfaces in a general scenario. But it can be easily shown that substitution using geometry parameters alone is possible for local spherical surfaces in contact, i.e., when k1m = k2m and k1f = k2f. In this case, the substitution-centers will be the centers of the contacting locally spherical profiles. A direct-contact mechanism consisting of an input link and an output link is shown in Fig. 10(a). Parts of the two links are torus surfaces which contact at point C on their respective outer circles. The input and output links are constrained to rotate about the zaxis and x-axis respectively. The minor radius and outer radius of the input link torus are 80 and 20 units respectively. Similarly those of output link torus are 15 and 90 units. The relative orientation angle is 30\u00b0. The input link is given angular velocity (\u03c9in) and angular acceleration (\u03bcin) and those of the output link (\u03c9out and \u03bcout) are to be determined for the configuration of themechanism shown in the figure. For input\u2013output kinematic relationship, substituted RUSR mechanism is equivalent to a RSSR mechanism. So, we use the RSSR mechanism output velocity and acceleration equations given in [22] for calculations. The reference ground points on the R-joints axes are chosen at (0, 0, 0) and (200, 0, 0). The point of contact is located at (100, 100, 50). In Fig. 10(b), the spherical joint locations for the substituted RSSRmechanism are chosen at (0, 200, 0) and (200, 0, 100). Velocity and acceleration analyses of this RSSRmechanism for chosen values of \u03c9in and \u03bcin give the following output: Input1: \u03c9in = \u22125 rad/s; \u03bc in = \u221210 rad/s2 Output1: \u03c9out = 10 rad/s; \u03bcout = 20 rad/s2. Now, we use the theory developed in Section 4 to obtain substituted RSSRmechanismswhich are valid up to acceleration analysis. For relativemotion between the two links, the input link is assumed to be fixed and a global fixed coordinate system is assigned at the point of contact as explained in Section 3 which is shown in Fig. 10(a). Twist coordinates of relative motion between the two links in this coordinate system are derived as: L M N P Q R 2 6666664 3 7777775 \u00bc ffiffiffi 2 p 6 \u03c9out \u00fe 2 ffiffiffi 2 p 3 \u03c9in 1ffiffiffi 2 p \u03c9out 2 3 \u03c9out\u2212 1 3 \u03c9in \u2212175 ffiffiffi 2 p 3 \u03c9out \u00fe 100 ffiffiffi 2 p 3 \u03c9in \u221225 ffiffiffi 2 p \u03c9out 200 3 \u03c9out \u00fe 400 3 \u03c9in 2 66666666666666664 3 77777777777777775 : \u00f064\u00de A MATLAB program is written (see Appendix A) to execute the velocity analysis, calculate the conjugate point coordinate zo for a chosen z value, and subsequently obtain the output angular acceleration",
            " We see that the output values\u03c9out and \u03bcout are same for both the above inputs. The substituted mechanisms for Input2 and Input3 are shown in Fig. 11(a) and (b) respectively. Now, we find output angular acceleration for Input2 taking a different zo other than the calculated value of\u2212 13.7794. If we take zo=\u221215, the output angular acceleration comes out to be \u03bcout=\u2212409.6875 rad/s2 which is different from the actual output value. Also, notice that \u03bcout for Input1 is different from that of Input2 and Input3; reason being nonconjugate pair of points chosen in Fig. 10(b). This shows that \u03bcout is the same for any conjugate pair of substitution centers but will cease to remain same if the points are chosen arbitrarily on the n-line. A binary link with U-joint and S-joint is substituted for the point contact between two surfaces to aid the instantaneous acceleration analysis of direct-contact mechanisms. Unlike the planar case, the substitution centers here cannot be determined by merely the curvatures at the point of contact. In addition, location of the ISA of relative motion between the surfaces and its pitch completely decides the conjugate point for a chosen point on the contact normal line"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000563_aim.2015.7222744-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000563_aim.2015.7222744-Figure3-1.png",
        "caption": "Fig. 3. Schematic representation of the local coordinate system",
        "texts": [
            " 2, the position vector r = [x, y, z] T of point O\u2032 can be expressed by r = bi + qie+ di + lwi \u2212 ai, (i = 1, 2, 3) (1) where ai = [ s cos\u03b2i \u2212s sin\u03b2i 0 ]T (2) bi = [ S cos\u03b2i \u2212S sin\u03b2i 0 ]T (3) di = [ \u2212d cos\u03b2i d sin\u03b2i 0 ]T (4) e = [ 0 0 1 ]T (5) where s, S, qi, \u03b2i, l, e, wi, di, ai and bi denote the radius of the moving platform, the radius of the fixed base, the linear displacement of slider, the angular of point Bi in the coordinate O \u2212 xyz, the length of the strut, the unit vector along the lead screw, the vector along the strut, the vector from a lead screw to the center point of universal joint Ci, the position vector of point Ai in the coordinate O\u2032 \u2212x\u2032y\u2032z\u2032 and the position vector of point Bi in the coordinate O\u2212xyz, respectively. Conventionally, the orientation of each strut with respect to the fixed base can be described by two independent Euler angles. As shown in Fig. 3, the local coordinate system of the ith strut can be imagined to be a rotation of \u03c6i about the zi axis resulting in a (x\u2032 i, y \u2032 i, z \u2032 i) system followed by another rotation of \u03d5i about the rotated y\u2032i axis. In this way, the rotation matrix of the ith strut can be expressed as oRi =   c\u03c6ic\u03d5i \u2212s\u03c6i c\u03c6is\u03d5i s\u03c6ic\u03d5i c\u03c6i s\u03c6is\u03d5i \u2212s\u03d5i 0 c\u03d5i   (6) where s\u2212 and c\u2212 denote sin and cos, respectively. Equating the third column of oRi to wi, yields wi =   c\u03c6is\u03d5i s\u03c6is\u03d5i c\u03d5i   (7) Solving Eq. (7) for \u03c6i and \u03d5i gives c\u03d5i = wiz s\u03d5i = \u221a w2 ix + w2 iy , (0 \u2264 \u03d5 \u2264 \u03c0) s\u03c6i = wiy/s\u03d5i c\u03c6i = wix/s\u03d5i (8) Taking the norm of both sides of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002423_jae-160067-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002423_jae-160067-Figure8-1.png",
        "caption": "Fig. 8. The experimental setup of deflection closed-loop control of cantilever beam.",
        "texts": [
            " As illustrated in Fig. 7(a), the deflection of cantilever beam takes about 41.7 seconds to increase to the target value when the PLZT ceramic voltage generator is illuminated by UV light of 200 mW/cm2 and the average fluctuation height f s 1 is about 0.4 \u03bcm. As illustrated in Fig. 7(b), the deflection of cantilever beam takes about 13.9 seconds to increase to the target value when the PLZT ceramic voltage generator is illuminated by UV light of 400 mW/cm2 and the average fluctuation height f s 2 is about 1.1 \u03bcm. Figure 8 shows the experiment setup of deflection closed-loop control of cantilever beam based on hybrid photovoltaic/piezoelectric actuation mechanism. The PLZT (3/52/48) patch is taken as the voltage source, which is provided by Shanghai Institute of Ceramics, Chinese Academy of Science. The PLZT specimens are electrically polarized in the air when the working temperature is above the Curie temperature. The dimensions of the PLZT specimen with polarization along the length direction are: 13 mm (length) \u00d7 5 mm (width) \u00d7 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002357_j.aej.2016.06.015-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002357_j.aej.2016.06.015-Figure2-1.png",
        "caption": "Figure 2 Single segm",
        "texts": [
            " (2) Bump stiffness is constant, independent of the actual bump deflection, not related to or constrained by adjacent bumps. (3) The top foil does not sag between adjacent bumps. The top foil does not have either bending or membrane stiffness, and its deflection follows that of the bump. With these considerations, the radial deformation of the bump foil due to the steady-state aerodynamic pressure (Wt) depends on the bump compliance (a) and the average pressure across the bearing width as shown in Fig. 2: Wt \u00bc K\u00f0p pa\u00de \u00f01\u00de where p and pa are the steady-state gas-film pressure and the ambient pressure, respectively and K is a constant reflecting the structure rigidity of the bumps. It was shown in [6] that K is given by K \u00bc aC pa \u00f02\u00de where a \u00bc 2paS CE l tb 3 \u00f01 t2\u00de \u00f03\u00de In order to study the performance of the foil bearing, we should first be able to select its dimensions. Eq. (3) shows that there are two main parameters that affect the compliance number, l the half length of bump in h direction and S the pitch of bump foil, so we should carefully select them",
            " The generalized Reynolds\u2019 equation is as follows: @ @x F2 ph3 l @p @x \u00fe @ @z F2 ph3 l @p @z \u00bc @ @x F4 F3 Fo u \u00fe @ @z F4 F3 Fo w \u00f04\u00de where Fo \u00bc Z h 0 dy l ; F1 \u00bc Z h 0 ydy l ; F2 \u00bc F1 Fo F3 Z h 0 q Z y 0 ydy l dy F3 \u00bc Z h 0 q Z y 0 ydy l dy;F4 \u00bc Z h 0 qdy If the variation in the lubricant\u2019s density is neglected, the generalized Reynolds equation in dimensionless form could be written in the following normalized form [20]: @ @h I2 P h3 @ P @h \u00fe D L 2 @ @ Z I2 P h3 @ P @ Z \u00bc K l @ @h P h 1 I1 Io \u00f05\u00de where Io \u00bc Z 1 0 d y l ; I1 \u00bc Z 1 0 yd y l ; I2 \u00bc Z 1 0 y l y I1 Io d y h \u00bc 1\u00fe e cos h\u00fe a\u00f0 P 1\u00de The appropriate boundary conditions for the Reynolds\u2019 equation are as follows: P \u00bc 1 at Z \u00bc 1 @ P @ Z \u00bc 0 at Z \u00bc 0 circulating air mixes with fresh air [18]. P\u00f0h \u00bc 0\u00de \u00bc P\u00f0h \u00bc hend\u00de where hend is the circumferential angle at which the top foil ends as shown in Fig. 2. Typically, hend \u00bc 355 [60]. The third boundary condition states that the pressure is periodic in the circumferential direction. Using the finite difference method to solve Eq. (5) and sim- plifying, the equation will be as follows: Table 1 Bearing data. Bearing radius, R= D/2 35 10 3 m Bearing length, L 70 10 3 m Bearing clearance, C 35 10 6 m Bump foil Young\u2019s modulus, Eb 207 109 N/m2 Bump foil Poisson\u2019s ratio, t 0.3 Table 2 Lubricant (air) data. Viscosity of air, la 1:932 10 5 Pa s Lubricant density, q 1:1614 kg/m3 Specific heat of air, Cp 1007 J/kg K Air conductivity, Ka 2:63 10 2 W/m K Shaft angular speed, x 30,000 rpm h3i;jI2 2 \u00f0Dh\u00de2 \u00fe D L 2 2 \u00f0D z\u00de2 \" # P2 i;j \u00fe h3i;jI2Pi\u00fe1;j h3i;jI2Pi 1;j \u00f0Dh\u00de2 D L 2 h3i;jI2Pi;j\u00fe1 \u00fe h3i;jI2Pi;j 1 \u00f0D z\u00de2 \u00fe K l 1 I1 Io hi\u00fe1;j hi 1;j 2Dh \" # Pi;j \u00fe h3i;jI2 Pi\u00fe1;j Pi 1;j 2 4\u00f0Dh\u00de2 D L 2 h3i;jI2 \u00f0Pi;j\u00fe1 Pi;j 1\u00de2 4\u00f0D z\u00de2 \u00fe K l 1 I1 Io hi;j Pi\u00fe1;j Pi 1;j 2Dh \" # \u00bc 0 \u00f06\u00de 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003257_j.ijnonlinmec.2020.103550-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003257_j.ijnonlinmec.2020.103550-Figure3-1.png",
        "caption": "Fig. 3. A Poincar\u00e9 map for fixed  = 0.1, \ud835\udc00 = diag(2, 3, 4), \u2126 = (\u22122, 0, 1), \ud835\udc53 = \u221a 5 and the secant \ud835\udc540 = 0. Red and blue denote separatrices. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " However, both in the case of existence of an additional integral (when the flow is restricted to the general integral surface) and in the case of existence of a symmetry field (when the flow is reduced to the space of orbits of a symmetry field), we obtain a two-dimensional autonomous system in which chaotic trajectories are known to be absent. Consequently, the existence of chaotic trajectories in this system leads to the absence of any of these invariants (both an integral and a symmetry field). We also note that in the system considered, at some parameter values, the chaotic trajectories group to form stochastic layers, see Fig. 3, and at other parameter values they form a strange attractor, see Fig. 6. We now illustrate the results obtained in the previous section by numerical analysis of the behavior of the trajectories of the system (1) on the integral surfaces 3 0,\ud835\udc53 (for which the Melnikov integral was calculated). For this we use a Poincar\u00e9 map. We first restrict the system (1) to the three-dimensional manifold 3 0,\ud835\udc53 using the Andoyer variables (\ud835\udc3f, \ud835\udc59, \ud835\udc54) [32]: \ud835\udc401 = \u221a \ud835\udc53 \u2212 \ud835\udc3f2 sin \ud835\udc59, \ud835\udc402 = \u221a \ud835\udc53 \u2212 \ud835\udc3f2 cos \ud835\udc59, \ud835\udc403 = \ud835\udc3f, \ud835\udefe1 = \ud835\udc3f \u221a \ud835\udc53 cos \ud835\udc54 sin \ud835\udc59 + sin \ud835\udc54 cos \ud835\udc59, \ud835\udefe2 = \ud835\udc3f \u221a \ud835\udc53 cos \ud835\udc54 cos \ud835\udc59 \u2212 sin \ud835\udc54 sin \ud835\udc59, \ud835\udefe3 = \u2212 \u221a 1 \u2212 \ud835\udc3f2 \ud835\udc53 cos \ud835\udc54, where \ud835\udc59, \ud835\udc54 \u2208 [0, 2\ud835\udf0b) are the angle variables and \ud835\udc3f, \ud835\udc53 satisfy the obvious inequality \u22121 \u2a7d \ud835\udc3f \u221a \ud835\udc53 \u2a7d 1",
            " We parameterize the manifold 2 \ud835\udc540 by a pair of variables \ud835\udc59 mod 2\ud835\udf0b and \ud835\udc3f \u221a \ud835\udc53 ( | | | | \ud835\udc3f \u221a \ud835\udc53 | | | | \u2a7d 1 ) so that the pair ( \ud835\udc59, \ud835\udc3f \u221a \ud835\udc53 ) defines a point on the two-dimensional unit sphere \ud835\udc462. A Poincar\u00e9 map for the Euler case ( = 0) is shown in Fig. 2. As can be seen, it is foliated by invariant curves, since the system has the additional integral (5) and the separatrices do not split. If condition (9) is satisfied, then the separatrices split and the separatrices from different periodic solutions intersect (see Fig. 3). Remark 2. If we set \ud835\udefa2 = 0, then the system (1) has the involution \ud835\udc61 \u2192 \u2212\ud835\udc61, \ud835\udc401 \u2192 \ud835\udc401, \ud835\udc402 \u2192 \u2212\ud835\udc402, \ud835\udc403 \u2192 \ud835\udc403, \ud835\udefe1 \u2192 \ud835\udefe1, \ud835\udefe2 \u2192 \u2212\ud835\udefe2, \ud835\udefe3 \u2192 \ud835\udefe3. In the Andoyer variables this involution is represented as \ud835\udc61 \u2192 \u2212\ud835\udc61, \ud835\udc3f \u2192 \ud835\udc3f, \ud835\udc59 \u2192 \ud835\udf0b \u2212 \ud835\udc59, \ud835\udc54 \u2192 \u2212\ud835\udc54 and generates the corresponding involution \ud835\udf0e of the Poincar\u00e9 map \ud835\udef1\ud835\udc53,\ud835\udc540 . As numerical experiments show, a large number of \u2018\u2018conservative\u2019\u2019 fixed points arise in this case on the Poincar\u00e9 map (i.e., the eigenvalues \ud835\udf061, \ud835\udf062 of the linearization matrix satisfy \ud835\udf061\ud835\udf062 = 1). As is well known, these \u2018\u2018conservative\u2019\u2019 points lie on the intersection of different iterations of involutions \ud835\udf0e\ud835\udc5b, \ud835\udc5b \u2208 Z, and these lines fill the Poincar\u00e9 map sufficiently densely (see, e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure11-1.png",
        "caption": "Figure 11. Four types of Bennett linkages (with a>b and j j 6\u00bc /2): (a) Type A(\u00fe) [b( ), a( )], (b) Type B(\u00fe) [b( ), a( )], (c) Type A( ) [b( ), a( )], and (d) Type B( ) [b( ), a( )].",
        "texts": [
            " The proper handedness of one chiral RR bar provides a privileged sense of path for going on Bennett\u2019s loop. The positive sense of path on the 4R loop can be selected as being the positive sense of path for going on the common perpendicular of an RR bar, which has properly a positive handedness. Using that convention, the signs of the acute angles in the RR bars are also those of the enantiomorphic types of bars. Summarizing the results, there are four RRRR Bennett linkages for a given system of two bar lengths, a and b, as well as two twist angles j j and j j. These RRRR chains described in Figure 11 can be derived from the four RRRS chains cited above when the four essential points of joints of RRRS make up an isogram. There are four possible sequences of oriented twist angles (or torsion angles) in the four bars: _ _ _ , _( )_ _( ), ( )_ _( )_ , and ( )_( )_( )_( ). They are sorted into two kinds: the kind A for _ _ _ and ( )_( )_( )_( ); the kind B for ( )_ _( )_ and _( )_ _( ). A plane symmetry changes any twist angle into its opposite. Hence, _ _ _ and ( )_( )_( )_( ) are the two enantiomorphic types (or two enantiomers) A(\u00fe) and A( ) of kind A; _( )_ _( ) and ( )_ _( )_ are the two enantiomorphic types B(\u00fe) and B( ) of kind B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure4-1.png",
        "caption": "Fig. 4. Detail view of fabricated the battery module",
        "texts": [
            " All the required mechanical elements, electrical PCBs, various sensors and odometers were integrated by the design shown in Fig. 3. The module was designed to expand/collapse a set of four drive-legs, and each wheel system contains a harmonic gear to rotate a drive-wheel using In-wheel motor. The battery module provides 25.2 VDC and up to 70 Ah. of energy to the robot for a meaningful 10 h mission. The battery module contains lithium polymer battery-cells combined into packs, safety electronics and voltage converters to allow monitoring of charge and discharge. An image of the layout of the module is shown in Fig. 4. The pump module was designed to provide the pneumatic actuator of the support system for the necessary centration. Also, it maintains 0.6 MPa of differential pressure to operate external pressure (5.0 MPa). DAS (Data Acquisition System) of NDE module, sub controller and hub PCB for communicating with NDE system and tractor robot were integrated (Fig. 5). The MFL system was designed to minimize friction between magnetizer and pipe wall, and maximize defect detectability. Also, the performance of the MFL system was verified, and it satisfies the specification of POF (Pipeline Operate Forum) standard"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003002_s00170-020-05276-z-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003002_s00170-020-05276-z-Figure16-1.png",
        "caption": "Fig. 16 Temperature of welded combustor casing",
        "texts": [
            " The results also affirm the superiority of the model\u2019s capacity to reflect real welding conditions and improve the accuracy of numerical calculations. According to the EBW heat source simulation and residual stress calculation, the low-cycle life prediction of welded combustor casing was also carried out in this study. A typical combustor casing from a civil aviation engine was used as an example (see Fig. 15). Like the welded plate used in Section 2 above, the thickness of the casing was 2 mm, so the EBW parameters remain unchanged. Given the operational load spectrum during actual flights, including the temperature (see Fig. 16), pressure, and axial force. The welding residual stress was added into the welded joint calculations in the form of a boundary condition, and its specific value was consistent with the calculated results in Section 3.1. TheManson\u2013Coffin [38\u201340] formula shown in Eq. (7) was used to predict the low-cycle fatigue (LCF) life of the welded combustor casing: \u03b5a \u00bc \u03c3 0 f 2Nf\u00f0 \u00deb=E \u00fe \u03b5 0 f 2Nf\u00f0 \u00dec \u00f07\u00de Here, \u03b5a is the strain range; Nf is the structural low-cycle fatigue life; \u03c3 0 f is the fatigue strength coefficient; \u03b5 0 f is the fatigue plasticity coefficient; b is the fatigue strength exponent; c is the fatigue plasticity exponent; and E is the elasticity modulus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002632_tencon.2016.7848445-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002632_tencon.2016.7848445-Figure1-1.png",
        "caption": "Fig. 1: Additive manufacturing process",
        "texts": [
            " Engineers, doctors, fashion designers, hobbyists are getting more freedom to innovate the design of their products. Conventional or subtractive manufacturing associated with design constraints, expensive tooling, fixtures, labor intensive and cooling systems which increases the cost of manufacturing. Researchers can overcome certain limitations associated with conventional manufacturing techniques using AM. AM is a process of obtaining objects by simply adding layer over layer of materials in 2D dimensions to get 3D parts(shown in Fig. 1). 3D printed parts can be complicate with minute design features and also capable of obtaining assembled final parts. AM is also considered as an sustainable manufacturing technique due to the low wastage of materials and energy requirements compared to conventional techniques[2][3]. In this paper application of 3D printing for improving the potential components of electrical power converters is elaborated. Power converters are essential part of electrical or electromechanical systems, application ranges from machine drive systems, renewable energy, electrical drive and More Electrical Aircrafts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002050_s10999-016-9338-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002050_s10999-016-9338-1-Figure2-1.png",
        "caption": "Fig. 2 Contact state and elastic\u2013plastic region distribution of asperity without friction: a asperity normal contact without friction and b distribution of elastic and plastic region (I elastic region, II plastic region)",
        "texts": [
            " dc refers to the critical deformation of the asperity varying from entirely elastic to elastic\u2013plastic stage. When d\\ dc, the asperities are in the range of elasticity. However, when d[ dc, the asperities are in elastic\u2013plastic stage. The contact stress is continuous in the interaction process of the two asperities, and the plastic deformation is generally generated in the adjacent region under the contact surface. Assuming that the normal distance between the initial plastic point O1 and the contact surface is y1 (see Fig. 2), according to JG model (Jackson and Green 2005) y1 is d dy rs F0 \u00bc ray1 r2 a 4 \u00fe m\u00f0 \u00de \u00fe 1 \u00fe m\u00f0 \u00dey2 1 \u00fe 1 \u00fe m\u00f0 \u00de r2 a \u00fe y2 1 2 tan 1 ra y1 \u00bc 0 \u00f011\u00de where F0 is the maximum normal contact load. Assuming that the contact region, whose distance is y1 from the contact point O, is composed of the plastic (inner circle) and elastic region (outside annulus), (see Fig. 2), and rp is the radius of the plastic region, the contact load of the region is written as FTp \u00bc pr2 prp \u00f012\u00de where rp \u00bc 2Et pR ffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 a r2 p q : From Eq. (9), the contact load of the elastic region is FTe \u00bc Z ra rp 2prr r\u00f0 \u00dedr \u00bc 4Et 3R r2 a r2 p 3 2 \u00f013\u00de The total contact load of the elastic and plastic contact regions is FT \u00bc FTe \u00fe FTp \u00bc 2Et 3R 2r2 a \u00fe r2 p r2 a r2 p 1 2 \u00f014\u00de Under the action of the normal contact load F, the plastic deformation occurs within the contact region of the radius rp, while the elastic deformation is still generated in other region",
            " Suppose that specific plane is across the initial point of plastic deformation and parallels to the bottom of the asperity, the contact stress isogram in the plane is plotted according to FE model. As shown in Fig. 10a, b, the contact stress isogram in the plane is plotted at the asperity normal or side contact state when the displacement load is 6dc. It is shown in the isograms that the contact stress gradually decreases from the center to around. In Fig. 10a, the contact stress isogram in the plane at asperity normal contact state is a circle, which is similar to the elastic\u2013plastic region distribution of the asperity without friction by the present theoretical model in Fig. 2b. Likewise, the isogram at asperity side contact state is a ellipse in Fig. 10b, which is similar to the elastic\u2013plastic region distribution of asperity with friction by the theoretical model in Fig. 3b. The comparative results show that the normal contact stresses of the asperities on the tooth profiles are similar according to the present elastic\u2013plastic asperity contact model with or without the consideration of friction, or to the finite element models for asperity normal and side contact analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure1-1.png",
        "caption": "Fig. 1 Planetary gear train with three external shafts, their torques and the lever analogy",
        "texts": [
            " It must not be forgotten that clarity is the foundation of engineering thinking and is a dominant requirement in engineering work. Unlike the methods of Willis and Kutzbach-Smirnov using angular velocity x (or rotational speed n) and the peripheral velocities v; at the presented here alternative method the torques T are used. The method is based on the following well-known principles of mechanics, illustrated with the simplest case of elementary (single-carrier) gear train with three external shafts that has a very good lever analogy (Fig. 1): 1. Equilibrium of ideal (with no account of losses) external torques acting on the three shafts X Ti \u00bc TD min \u00fe TD max \u00fe TR \u00bc 0 \u00f01\u00de where TD min and TD max are the unidirectional external ideal torques (with different sizes), therefore are indicated by single lines on the figure, but with different thickness, according to their size; TR is the largest torque, in absolute value equal to the sum of other two torques TD min and TD max; and therefore is marked with a double line. 2. Equilibrium of the real (considering the losses) external torques X T 0i \u00bc T 0D min \u00fe T 0D max \u00fe T 0R \u00bc 0 \u00f02\u00de where T 0D min; T 0 D max and T 0R are the same external torques as above, but real now considering the losses in the gear train",
            " not with numbers, letters or inscriptions, but by the thickness of the lines corresponding to the size of the external torque, as already mentioned above. This little innovation concerning the manner of marking the shafts makes the method very clear and useful. 6. Besides using the modified symbol of Wolf, where appropriate another innovation is made. A new torque ratio is defined\u2014the one of the unidirectional external torques (Arnaudov 1984, 1996; Arnaudov and Karaivanov 2001, 2005a, b; Karaivanov 2000). t \u00bc TD max TD min [ \u00fe 1 \u00f07\u00de because TD max [ TD min (Fig. 1). This innovation is especially important. It is connected to the lever analogy. 7. Ideal external torques are always in a certain ratio, expressed with the torque ratio t TD min : TD max : TR \u00bc T1 : \u00fet T1 : 1\u00fe t\u00f0 \u00deT1 \u00bc 1: \u00fet : 1\u00fe t\u00f0 \u00de \u00f08\u00de in which always TD min\\TD max\\ TRj j \u00f09\u00de however: \u2022 How many degrees of freedom F (F = 1 or F = 2) the gear train is running with; \u2022 Which shaft is fixed at F = 1 degree of freedom; \u2022 What is the direction of power flow, respectively does the gear train works as a reducer or a multiplier at F = 1 degree of freedom, respectively such as a collecting or as a separating gear at F = 2 degrees of freedom, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001567_c5sm01956g-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001567_c5sm01956g-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the experimental configuration for observing the electric field response of particles.",
        "texts": [
            " A square wave electric field with a frequency of 10 kHz was applied between the interdigitated electrodes and the response of the particles was recorded using a digital camera. The dynamic response of the particles was measured by inserting a band-pass filter (l = 550 nm, Dl = 40 nm) in the optical path and measuring the change in output light intensity using a photomultiplier tube. A 200 mm-core optical fiber was used in conjunction with a 100 objective lens, corresponding to a spot diameter of approximately 2 mm. Fig. 1 depicts the experimental configuration. Square-shaped particles with interior molecular alignment along the diagonal are floating in a sandwich cell with unidirectional LC alignment. The direction of the alignment is at 451 to the electrodes. Fig. 2 shows the POM images of a 20 mm particle dispersed in the 5CB host. The direct laser writing process enables particles with uniform interior orientations to be fabricated, which allows us to obtain colloidal dispersions without any topological defects"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.78-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.78-1.png",
        "caption": "Figure 4.78 Linear motion; belt and pulley drive",
        "texts": [
            "8)) = +1246 g cm TRUN = 1805 g cm TSS = 0 242 Electromechanical Motion Systems: Design and Simulation In order to determine TRMS, the run time must be calculated from the specifications that is, total travel is 107 cm, run velocity is 12.7 cm s\u22121 and acceleration and deceleration times are 0.2 s, therefore: (0.5) (0.2) (12.7) (2)+ (12.7) (tRUN) = 107 tRUN = 8.2 s TRMS = \u221a 23632 (0.2)+ 18052 (8.2)+ 12462 (0.2) 0.2+ 8.2+ 0.2 = 1809 g cm IACC = 2363/1236 = 2 A IDEC = 1246/1236 = 1.0 A IRUN = 1805/1236 = 1.5 A IRMS = 1809/1236 = 1.5 A Similar to the belt and pulley drive for rotary load motion, the inertias of the pulleys and the weight of the belt must be considered in determining the total motor load (see Figure 4.78). System Data: JM = motor inertia (g cm s2) JC = coupling inertia (g cm s2) W = load weight (g) FL = load force (g) FG = gravity force = W (g) FF = friction force = \u03bcW (g) BW = belt weight(g) JPM = motor pulley inertia (g cm s2) System Design 243 JPI = idler pulley inertia (g cm s2) e = screw efficiency \u03bc = friction coefficient Motion: Position: \u03b8M = S DPM/2 (rad) Velocity: \u03b8 \u2032 M = S\u2032 DPM/2 (rad s\u22121) Acc/Dec: \u03b8 \u2032\u2032 M = S\u2032\u2032 DPM/2 (rad s\u22122) Belt weight reflected to motor as an equivalent inertia: JB = ( BW 980"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000716_ecce.2015.7309909-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000716_ecce.2015.7309909-Figure7-1.png",
        "caption": "Fig. 7: Flux density map (a) SPM machine and (b) REL machine with eccentricity on the right hand side by e = 0.2 mm at \u03b8e = 0\u25e6, rotor position \u03b8m = 0\u25e6.",
        "texts": [
            " The location of the Q-axis depends on the type of eccentricity as shown in Fig. 1 (b) and Fig. 1 (c). In the following comparison, both SPM machine and REL machine have the same geometrical dimensions. The main data used in the analytical and FE simulations are listed in Table I. A mechanical air gap length of 0.35 mm is considered. In the SPM machine, a bandage thickness 0.2 mm is added so that the total magnetic air gap becomes equal to 0.5 mm. The 2D FE model is used to validate the results achieved by the analytical model. Fig. 7 shows both the flux lines and the flux density map of the SPM machine and REL machine with eccentricity e = 0.2 mm at \u03b8e = 0\u25e6 when the rotor is at the position \u03b8m = 0\u25e6. A displacement e = 0.2 mm is quite high, however it allows the eccentricity effect to be better highlighted. For both machines, the flux density increases on the right hand side where the air gap length is reduced and decreases on the left hand side where the air gap length is increased, respectively. It is worth noting that in the SPM machine the flux lines remain quite symmetrical and the flux density quite similar among the poles",
            " The main reason is the presence of the surface PMs which tend to keep a constant flux density even with the air gap variation. On the contrary, in the REL machine the flux lines are quite distorted and the flux density increases greatly near the lower air gap length. This is mainly due to the low distance between the stator and the rotor iron. Fig. 8 shows the air gap flux density of SPM and REL machines with and without rotor eccentricity. The flux density behaviors in case of eccentricity correspond to the maps in Fig. 7. With uniform air gap, the flux density is the same in all four poles, while it is not in case of eccentricity. It increases where the air gap length is reduced and decreases where the air gap length is increased. Fig. 8 confirms that the higher impact is on the REL machine where the variation of flux density is greater when eccentricity occurs. Fig. 9 shows the air gap flux density computed from the analytical and FE models of SPM and REL machine with eccentricity equal to 0.2 mm. It is noted that there is a satisfactory agreement between the analytical models and the FE analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000573_acc.2015.7171118-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000573_acc.2015.7171118-Figure1-1.png",
        "caption": "Fig. 1: Wind Turbine with Gearbox",
        "texts": [
            " As with the two mass model, the model analysis proceeds with the referral of all physical quantities to the low speed shaft. There are examples of the three mass model in the literature [8-9]. The five mass model consists of a generator inertia, rotor inertia, and lumped inertias for the planetary drive and two parallel drive stages. The planetary and parallel stages representation is a common gearbox configuration, but other gearbox configurations can also be built. In the five mass model, there are four connecting shafts for which stiffness and viscous damping can be specified. Fig. 1 shows the configuration of the five mass model. The wind turbine literature includes several examples of the gearbox types that can be represented by the five mass model usage [10-13]. In two of these cited examples [10-11], the drive train was represented by the lumped parameter model. However, two of the cited examples [12-13] used the MATLAB Toolbox Simscape from Mathworks [14] to represent individual gears, gear tooth stiffness, gear torsional stiffness, and shaft stiffness. The degrees of freedom included in the Simscape model are greater than that of the five mass lumped parameter degrees of freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002408_cdc.2016.7799264-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002408_cdc.2016.7799264-Figure2-1.png",
        "caption": "Fig. 2. 3-DOF helicopter model",
        "texts": [
            " Finally, simulation results are presented in Section VI. Conclusions are drawn in Section VII. The considered system (see Figure 1) is the 3-DOF Quanser helicopter [6], which is widely proposed in literature for experimental setup of control approaches. This system represent an underactuated 3-DOF mechanical system, driven 978-1-5090-1837-6/16/$31.00 \u00a92016 IEEE 6464 by two DC motors. All the parts are connected with revolute joints and the following three variables are analyzed: the elevation angle \u03b5, the pitch angle p and the travel angle \u03bb. See Figure 2 for the considered parameters and variables. This setup consists of a base on which a long arm is mounted. The arm carries the helicopter body on one end and a counterweight on the other end. The arm can tilt on an elevation axis as well as swivel on a vertical (travel) axis. Quadrature optical encoders mounted on these axes measure the elevation and travel of the arm. The helicopter body, which is mounted at the end of the arm, is free to pitch about the pitch axis. The pitch angle is measured via another encoder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000288_humanoids43949.2019.9035051-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000288_humanoids43949.2019.9035051-Figure1-1.png",
        "caption": "Fig. 1. The figure illustrates the experimental setup used for IMUbased data collection. A dynamometer was used to measure the contact force applied by the finger, while the IMUs were used to measure the Metacarpophalangeal (MCP) and Proximal Interphalangeal (PIP) joints. This experiment was conducted for 10 different finger angles from 0\u25e6 to 90\u25e6 in 10\u25e6 steps. The motor, the finger and the dynamometer were connected to bases that were fixed to an acrylic plate. The experiments were repeated for 10 trials for every experimental condition.",
        "texts": [
            " The finger structure consists of a combination of Polylactic Acid (PLA) plastic links and a polyurethane elastomer (urethane rubber Smooth On PMC-780) that is used on the fingerpads to increase the friction between the finger and the object during object handling, while the elastomer at the distal end of the finger acts like a distal interphalangeal joint (DIP) that has limited mobility and offers only conformability to the object surface. The robotic finger was designed with a hlwell-known, tendon-driven actuation system for adaptive and underactuated grippers [25]\u2013[27]. A. IMU-based Joint Angles Data Collection The finger was mounted onto a test structure with detachable mounts that allows us to vary the distance between the finger and the sensor as well as the angular parameters. The experimental setup is shown in Fig. 1 with the dynamometer mounted with a 30\u25e6 angle on the acrylic plate. Customized, non-conductive mounts were prepared for the IMU sensors (MPU-9250) to be attached onto the finger structure without interfering with the fingertips and the phalanges reconfiguration. The Biopac MP36 data acquisition unit (Biopac Systems, Inc., Goleta, California) was equipped with the SS25LA dynamometer and it was used to collect the contact force measurements. The motor (Dynamixel XM430-W350-R) and the finger were mounted on a custom acrylic plate while the dynamometer was attached onto offset acrylic plates that had an angle different between 0\u25e6 and 90\u25e6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002820_j.jcsr.2020.105959-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002820_j.jcsr.2020.105959-Figure4-1.png",
        "caption": "Fig. 4. Loading program. Fig. 5. Stress distribution for plate at \u0394y.",
        "texts": [
            " Three linear variable differential transformers (LVDTs) were equipped for measuring the displacement at points 1\u20133 in order to calculate the joint rotation by formulas (1) and (2). Fig. 3(b) shows the specimen during the test under loading. An apparent deformation appeared on the side plates of the specimen as the load increased. \u03a6 \u00bc P \u03d5ij 3 \u00f01\u00de \u03d5ij \u00bc arctan \u03b4i\u2212\u03b4 j lij \u00f02\u00de \u03b4i and \u03b4j, are the displacement values measured by the LVDTs, and lij is the distance between points i and j on the specimen. The displacement-controlmodewas adopted in this study, as shown in Fig. 4. \u0394y is the displacement value of the end plate as the edge of the side plate yields (Fig. 5), whichwas obtained using FEA in [18]. Each cyclic reverse loading was repeated twice. The specimens were tested until a large plastic deformation occurred or the carrying capacity decreased abruptly. joint. The hysteretic curves of the joints under cyclic loads can reflect the mechanical performance of the joints, including the strength, stiffness, ductility, and energy consumption. The M-\u0424 hysteretic curves and failure modes are shown in Figs",
            " The fixed constraint is set in the column node. The contact settings are shown in Fig. 12 aswell: the tie contact is used to constrain two separate surfaces together so that there is no relative motion between them. A surface-to-surface contact with a friction coefficient of 0.4 is set in the joint model between the column node and front plate, front plate and bolt, etc. A horizontal reciprocating load is applied to the upper end of the member with the displacement control method, which is similar to the test loading in Fig. 4. The pretension value in the high-strength bolts is 225 kN. The stress-strain relation of each component is determined according to a property test [18]. The specific material properties of each part are listed in Table 2. 3.2. Verification of FEM The numerical and test results including the hysteretic curves, skeleton curves, and failure mode are shown in Figs. 13\u201315. The hysteretic curves obtained experimentally are consistent with the numerical ones, thus experimentally verifying the effectiveness of the FE models"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure19-1.png",
        "caption": "Figure 19. Forces on a flapping airfoil. (a) A bird\u2019s wing moves down and forward on the downstroke, and this tilts the lift vector forward to produce thrust as a horizontal component. (b) Upstroke and downstroke of a flapping hydrofoil (porpoise\u2019s tail), where the horizontal component of the lift, Lh, represents thrust, and the vertical components of lift, Lv, on the upstroke and downstroke cancel. L: lift; Lu: upward component of lift; T: thrust; \u03b1: angle of attack.",
        "texts": [
            " These birds all have high lift-to-drag ratios, either due to high aspect ratios or due to modified wingtips that give some of the effects of high aspect ratios (slots between tip feathers, \u2018separated primaries\u2019; Alexander 2002). If an animal flies level continuously or climbs, it needs to have some source of thrust, a forward force to overcome drag. Flying animals flap their wings to produce thrust, and the most important part of the flapping stroke is the downstroke. During the downstroke, a flapping wing moves down and forward, at some positive angle of attack. As a result of the downward movement, the lift vector tilts forward (figure 19(a)). The upward component of lift supports the flyer\u2019s weight. The lift also has a horizontal component directed forward, and this horizontal component is the thrust. Thrust production is the function of flapping, and if thrust equals total drag, then the flyer will be able to maintain level flight, thus overcoming the downward requirement of gliding. Increasing the thrust allows climbing, and decreasing the thrust allows descent. Upstrokes can vary greatly, depending on the type of animal. They can be a largely passive recovery stroke, with angle of attack and area adjusted to minimize forces, as in most medium and large birds",
            " These swimmers are very near the same density as the water they inhabit, so little or none of their locomotory effort needs to go into supporting their weight. Consider the tail of a cetacean such as a porpoise. It flaps up and down like a bird wing, but since the porpoise needs no net upward force to support its weight, the downstroke occurs at a positive angle of attack and generates lift upward and forward, with a horizontal component providing thrust. The upstroke uses a negative angle of attack, so the \u2018negative\u2019 lift vector points down and forward, forming an upside down mirror image of the downstroke (figure 19(b)). The vertical components (Lv) of the upstroke and downstroke cancel, and the horizontal components (Lh) provide an near-continuous forward thrust. The stroke can be tweaked a bit to give a bit of net downward force for positively buoyant animals \u2014those slightly less denser than water, like penguins at shallow depths\u2014or a bit of net upward force for negatively buoyant animals\u2014like whales and seals at great depths\u2014but the vertical forces involved are tiny compared to the horizontal forces of swimming"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002476_tac.2016.2631302-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002476_tac.2016.2631302-Figure1-1.png",
        "caption": "Fig. 1: (a) Cyclic pursuit scheme. (b) Eigenvalues.",
        "texts": [
            " Consensus in the asynchronous mode is also achieved with heterogeneous gains, when all the gains are positive. D. Mukherjee is currently a Postdoctoral fellow at the Faculty of Aerospace Engineering, Technion- Israel Institute of Technology, Haifa, Israel, e-mail: dwaipayan.mukherjee2@gmail.com. D. Ghose is a Professor at the GCDSL, Department of Aerospace Engineering, IISc, Bangalore, India, e-mail: dghose@aero.iisc.ernet.in. This work was partially supported by NP-MicAV & AOARD. The cyclic pursuit topology, with n agents, is shown in Fig. 1a. Each agent (say agent i) follows its leader (agent i + 1 modulo n) and decides its position at the next sampling instant based on the length and direction of the vector joining it to its leader. A synchronous system, where the agents update their positions simultaneously, is assumed. For conciseness of notation, define zi(t) as a complex number that represents the position of agent i at time instant t. The position vector of the agents is expressed as z(t) = [z1(t) . . . zn(t)]T \u2208 Cn. In discrete cyclic pursuit, the update rule for agent i is given by zi(t+ 1) = (1\u2212\u03c1i)zi(t) +\u03c1izi+1(t), where the indices of the agents are modulo n",
            " The system will converge to the space spanned by the eigenvector corresponding to simple eigenvalue 1 if all the other roots of A lie within the disc of unit radius, centred at the origin of the complex plane [9], [10]. The characteristic equation of A is (s \u2212 1 + \u03c1)n \u2212 \u03c1n = 0, whose roots, pr = 1 \u2212 \u03c1 + \u03c1ej 2\u03c0 n r, r = 0, 1, . . . , n \u2212 1, are evenly spaced along the periphery of the circle with its centre 0018-9286 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. at (1 \u2212 \u03c1, 0) and radius \u03c1 (Fig. 1b). This follows from the property of circulant matrices [11]. If 0 < \u03c1 < 1 then all the eigenvalues, except the one at (1 + j0), are inside the unit circle, shown in Fig. 1b. This guarantees positional consensus of the system. This result is identical to the one obtained in [2] which uses Gershgorin\u2019s Theorem to establish stability for 0 < \u03c1 < 1. Here, the eigenvalues are explicitly given. The heterogeneous case is now considered. Lemma 1. Consider the system described by (1)-(2). If 0 < \u03c1i < 1, \u2200i, then the system achieves positional consensus. Proof. As the union of the Gershgorin discs lies inside the open unit circle in Fig. 2a (dotted circle), except at the point (1, 0) where the Gershgorin discs are tangential to the unit circle, it suffices to ensure that there are no repeated roots at 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure18-1.png",
        "caption": "Fig. 18. The rotor temperature distribution diagram when the rotor loss is 250W. (a) original structure, (b) optimized structure with blades.",
        "texts": [
            " 2 uses eight heating pipes evenly distributed on the outer edge of the rotor, which is used to equivalent the actual rotor loss. In order to highlight the optimization of the cooling effect of the recirculating hollow-shaft, the rotor experimental platform was modeled and the temperature field simulation was performed using Workbench software. It can be seen from Fig. 17 that the rotor average temperature of the optimized structure is lower than that of the original structure, and the temperature difference between the two increases as the rotor loss increases. Fig. 18 shows that at a rotor loss of 250W, the temperature of the optimized structure is 33.6oC lower than the original structure. 3516 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSIONS This paper has evaluated and compared the CHTC and friction loss of recirculating hollow-shaft and direct-through hollow-shaft. It shows that the recirculating hollow-shaft is superior to the direct-through hollow-shaft in terms of the CHTC, but friction loss of direct-through hollow-shaft is lower"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000741_jfm.2014.45-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000741_jfm.2014.45-Figure4-1.png",
        "caption": "FIGURE 4. (Colour online) The sketch of a goose and the wing-tip vortex model, where \u0393 is the strength of the vortex, u is the downwash velocity induced by the wing circulation and the wing-tip vortices themselves and \u0393 \u00d7 u is the Lamb vector.",
        "texts": [
            "2) where b is the width between wing tips and \u03b4 is a desingularization factor which can be regarded as the dimensionless core radius. The distribution of the downwash velocity distribution along the horseshoe vortex leg is shown in figure 3, where \u03b4 = 0.15. It is found that there is a region of large induced velocity near the wing tip (x/b < 1), i.e. the downwash velocity of the wing-tip vortex is large when it is near the wing and has a characteristic of rapid decrease to a constant value with the increase of longitudinal distance from the wing tip. A sketch of a goose and the wing-tip vortices is shown in figure 4. The strength of the vortex is \u0393 , and the downwash velocity induced by the wing circulation and the wing-tip vortices is u. Then, the Lamb vector in the wake of the goose is \u0393 \u00d7 u, whose magnitude distributed along the longitudinal direction has the same feature as the induced downwash velocity. In addition, the Lamb vector directs outward along the span of the wing. According to the virtual power principle, the peak of the Lamb vector in the near wake of the wing tip could benefit the flight of the other geese"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000594_icarsc.2015.38-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000594_icarsc.2015.38-Figure3-1.png",
        "caption": "Fig. 3: Forward motion analysis when rotors rotate with standard quadrotor configuration",
        "texts": [
            " But the possibility of tilting two of its rotors also raises some performance problems when compared to a common quadrotor. The presence of two swivelled motor-propeller sets, each with two tilting angles, introduces a substantial performance change with respect to a common quadrotor. While in a common quadrotor the pair of rotors (1,3) rotates in a clockwise direction and the pair (2,4) rotates in a counter-clockwise direction in order to annul the resulting moment, for the tiltquadrotor this balance only exists in a hover situation. Consider the forward motion represented in fig. 3, where rotors 2 and 4 have, respectively, pitch-tilt angles \u03b82 and \u03b84 (positive angles are defined in order to induce a positive yaw). If both opposing rotors rotate in the same direction, the resulting moment in the horizontal plane is not annulled, inducing an undesired roll motion in the tilt-quadrotor. Bearing in mind the different possible configurations of the tilt-quadrotor, a performance analysis was made that lead to the following conclusions: \u2022 The rotors pairs 1-4 and 2-3 must rotate in opposing directions, respectively counter-clockwise and clockwise directions, in order to cancel the undesired moments induced by the rotors tilting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002363_0142331216661760-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002363_0142331216661760-Figure2-1.png",
        "caption": "Figure 2. Illustration of the main idea of blade element theory (BET) on a propeller blade (left). During flight simulation, the aircraft is split into a number of surfaces and the forces on each are computed by BET (right).",
        "texts": [
            " The coefficients can be selected empirically or via automated tuning methods such as Ziegler\u2013Nichols or internal model control (IMC) (Ho et al., 1995). This finalizes the construction of the flight control system. The control system is verified with software-in-the-loop (SIL) simulations incorporating BET. The surfaces of the aircraft (e.g. propellers, wings and stabilizers) are divided into several sections, the lift/drag forces acting on each section are computed separately and the composite effect is applied to the entire aerial vehicle (Figure 2). This approach contrasts with traditional flight simulations relying on empirical data (e.g. stability derivatives) in predefined lookup tables and is widely accepted to be more realistic albeit computationally expensive. It is also a good choice for testing the control design presented here, as the mathematical model utilized is based on stability derivatives. Hence, a different (and more accurate) flight simulation technique that does not rely on stability derivatives serves as a better test. The main idea of BET can be at CORNELL UNIV on September 6, 2016tim.sagepub.comDownloaded from summarized on a propeller blade shown in Figure 2 (left). The blade is divided into N elements, each of which experiences a slightly different flow. Lift and drag coefficients (CL and CD) are readily available for numerous airfoil shapes from wind tunnel tests. Using relative velocities, the flow over each element can be related to these tests. The flow is slightly turned passing over the airfoil so inlet and exit flow conditions are averaged to improve accuracy. Carrying out the necessary computations yields dFx = dL sinb+ dD cosb \u00f038\u00de dFu = dL cosb dD sinb \u00f039\u00de dL=s0pr V 2 1 a\u00f0 \u00de2 cos2b CLrdr \u00f040\u00de dD=s0pr V 2 1 a\u00f0 \u00de2 cos2b CDrdr \u00f041\u00de s0= Bc 2pr \u00f042\u00de where dL, dD, dFx and dFu are respectively the lift, drag, axial and tangential forces, b is the relative flow angle, r is the air density, V is the flow velocity, a is the axial induction factor, r is the radius, s0 is the local solidity, B is the number of blades and c is chord length (Ingram, 2005). The procedure is carried out on the entire aircraft to compute all the forces, using which the flight dynamics can be simulated as in Figure 2 (right). This step concludes the design and verification of the controller design. A flowchart of the entire process is given in Figure 3. A case study is presented in the next section. In this section, the autopilot design methodology outlined in the previous section is demonstrated on a popular general aviation aircraft, namely the Cessna 172. The geometry and mass parameters of this aircraft, together with its stability derivatives and performance specifications are given in Table 1. Based on the table, a 63 6 grid of operating points are constructed with all possible combinations of airspeeds v= 25 35 45 55 65 75\u00bd and elevation values ze = 0 500 1000 1500 2000 2500\u00bd "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003409_j.engfailanal.2020.104863-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003409_j.engfailanal.2020.104863-Figure4-1.png",
        "caption": "Fig. 4. Finite element model for stick fatigue bench test.",
        "texts": [
            " In other words, both elements are well-suited for linear applications. Furthermore, each bolt was set up by nodal degree of freedoms (DOFs) coupling, such that the X, Y and Z direction displacement of the two nodes on the bolt were consistent. Rigid constraints were used to simulate the axis pin connection between the stick and test bench. The displacement of lower beam and ground were constrained. The total number of elements and nodes were 27,071 and 81,702, respectively. The FEM of stick fatigue bench test is shown in Fig. 4. The maximum value of the test load spectrum was 263,680 N, which was applied on the front-end of stick, and the load direction is the positive direction of Y. Mechanical properties of bolts are listed in Table 3. Considering the load case, assembling conditions, materials and structure of bolts connection, allowable tensile stress [\u01a1] was taken from the specification. The pulling force (FY) of each bolt was determined using FEM. Then, the tensile stress value ca of each bolt was obtained. The distribution of tensile stress for bolts connection is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.17-1.png",
        "caption": "FIGURE 14.17 Pole structure\u2014section at height z.",
        "texts": [
            " NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION the pole is to be installed, a topographic factor that depends on the surrounding terrain, and an importance factor for the structure, as given by the Nbr-6123 code (AbNT, 1988). The effective wind velocity (V0) is also called the maximum wind velocity, distinguished from Voper (operational wind velocity). Therefore, q(z) in Eq. (14.153a) is also called the maximum wind pressure. The internal loads, stresses, displacements, rotations, and curvatures derived from q(z) are based on the maximum wind velocity. The expressions for the loads and moment in Eqs. (14.153h)\u2013(14.153l) are derived by considering a section at height z, as shown in Fig. 14.17. Note that the vertical load expression in Eq. (14.153i) contains the self-weight of the pole. This implies that total axial load and moment at a point are dependent on the pole\u2019s design. To develop expressions for the cost and constraint functions, considerable background information is needed. First, we need to calculate stresses in the pole section at height z, which requires analysis of the structure. The members are subjected to both bending moment and axial load, and the stresses of these internal loads are combined in Eq",
            " The lagrange multiplier for the tip deflection constraint is 673. N(0)A(0)+M(0)S(0)\u2264\u03c3a v(H)\u2264va. 0.30\u2264dt\u22641.0 m 0.0032\u2264t\u22640.0254 m 0\u2264\u03c4\u22640.05 m/m \u03c3\u03c3a\u22641 II. NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION EXAMPLE 14.11 OPTIMUM DESIGN WITH THE TIP ROTATION CONSTRAINT In practice, the pole\u2019s antennas must not lose the link with the receiver under operational wind conditions due to wind velocity, Voper. In this case, the rotation of the antennas must be smaller than a given limit, called the maximum rotation allowable for the antennas \u03c5\u2032( )a . Fig. 14.17 shows this rotational limit constraint. Antenna A is installed on pole A and has a link with antenna b on pole b. because the antennas are fixed on the pole, if the poles rotate more than the allowable value for them, they will lose their link and the system will go off the air (Fig. 14.18). For that reason, we impose a new constraint that is related to tip rotation, which must be within its allowable value: \u2032 \u2264 \u2032v H v( ) aoper (14.162) Note that to obtain \u2032voper, it is necessary to integrate the elastic line equation, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002968_lra.2020.2986746-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002968_lra.2020.2986746-Figure2-1.png",
        "caption": "Fig. 2. Fabrication of a single sensor: (a) 15 g of Ecoflex 00\u201310 mixture is poured into a mold to form a layer of 1 mm thickness. (b) The cured Ecoflex is covered with a mask and the Carbon Black composite is spread across the channel. (c) The mask is peeled off leaving a conductive sensor channel and a second layer of 10 g of Ecoflex 00\u201310 mixture is poured. (d) The membrane is slowly peeled from the mold. (e) The carbon composite is composed of a 1:6 mass ratio of Carbon Black powder to Ecoflex 00\u201330.",
        "texts": [
            " We increased the carbon black to Ecoflex ratio to 1:6 and mixed it by hand for several minutes. Doing this resulted in an uncured conductive gel that can be spread onto a polymer surface in a thin layer, as discussed in Sec. II-B. The cost of making a quarter coin-sized conductive composite with this technique is only 0.12 USD, saving an additional 18% compared to the pilot experiment. This simplified conductive composite fabrication technique requires no specialized equipment, takes only a few minutes, and is orders of magnitude faster than other common approaches. Fig. 2 shows the process for embedding the carbon composite in Ecoflex gel to fabricate single sensors. First, we pour a 1 mm thick base layer Ecoflex rubber mixed 1A:1B by weight into a mold. To eliminate entrapped air, we vacuum degas the mixture before pouring. We cure the silicone for approximately 2 hours at room temperature, or speed up the process by using an oven at 70\u00b0 C. After the silicone has cured, we cover it with a laser-cut stencil which is 0.8 mm in thickness. The stencil is 2 mm wide and corresponds to the width of the sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001323_1.4024555-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001323_1.4024555-Figure6-1.png",
        "caption": "Fig. 6 HDI 1.6 L 110 CV connecting rod",
        "texts": [
            " Hence, several responses have to be monitored: for the bearing performance (e.g. the power loss), but also for the bearing durability (e.g. the minimum of the film thickness). The input parameters do not have the same advantageous effects on the aforementioned responses. As stated above, one can reduce the power loss while critically decreasing the minimum of the film thickness: one speaks of competing objectives. 6.1 The Connecting Rod Big-End Bearing. The studied bearing is the connecting rod big-end bearing of a European HDI 1.6L 110 CV engine (Fig. 6). The engine regime chosen is 2000 rpm at full load (Fig. 7). It is a harsh case, for which facing very nonlinear behaviors of the responses regarding the input parameters can be expected. However, as the goal of this paper is to prove the efficiency of a methodology based on metamodels, this case is relevant for a global validation. To complete the connecting rod description, Table 1 gives additional information. 6.2 The Input Parameter Inventory. Following SAE (Society of Automotive Engineers) standard, the rheological properties of a 5 W/30 oil are given within a range: for the grade \u201c30,\u201d the kinematic viscosity is comprised between 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001610_0959651814520827-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001610_0959651814520827-Figure6-1.png",
        "caption": "Figure 6. Experimental bench of APS.",
        "texts": [
            " e d NB NM NS ZO PS PM PB kp/ki/kd NB: negative big; NM: negative middle; NS: negative small; ZO: zero; PS: positive small; PM: positive middle; PB: positive big; SS: small; SM: middle small; MM: middle; BM: middle big; BB: big. Figure 5. Simulation results of APS. The dashed-dotted line shows the reference signal, the solid line shows the response of APS and the dashed line shows the tracking error. at RYERSON UNIV on June 16, 2015pii.sagepub.comDownloaded from Experiments have been conducted to demonstrate the performance of the proposed controller. Figure 6 is the photo of the experimental system. The pneumatic system includes a vacuum pump (Edwards nXDS15i) and a 5-way servo valve (Norgren VP60) used to regulate the absolute pressure of the chamber (FESTO-80-0.1SAS073). The pressure sensor (Setra 270-RoHS) as well as the computer control system with 16-bit Advantech high-speed A/D D/A card (PCI 1716) are used to get the system output and drive the servo valve to change air-flow rates into and out of the closed chamber. Figure 7 describes the experimental results with conventional PID controller and feedback linearizationbased self-tuning fuzzy PID controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000323_humanoids43949.2019.9336618-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000323_humanoids43949.2019.9336618-Figure3-1.png",
        "caption": "Fig. 3: (a) Kinematic model of Cassie showing the robot\u2019s generalized coordinates in the body frame. (b) Hovershoe model where u\u03b8 and u\u03c8 change the pitch and yaw of the Hovershoe in the X-Y-Z body frame, respectively.",
        "texts": [
            " Cassie has twenty DOFs as listed in (1): q = [qx, qy, qz, qyaw, qpitch, qroll, q1L, q2L, q3L, q4L, q5L, q6L, q7L, q1R, q2R, q3R, q4R, q5R, q6R, q7R] T , (1) where, (qx qy qz) and (qyaw qpitch qroll) are the Cartesian coordinates of the pelvis and the Euler Angles in the Z-YX order, and (q1L, ...,q7L), (q1R, ...,q7R) are the generalized coordinates of the left and right legs, respectively. These correspond to the DOFs for each leg and are defined in (2).\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 q1 q2 q3 q4 q5 q6 q7 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 hip roll hip yaw hip pitch knee pitch shin pitch tarsus pitch toe pitch \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 . (2) Fig. 3a shows the generalized coordinates of Cassie\u2019s pelvis and right leg. The generalized coordinates of Cassie\u2019s left leg are similar to the right leg states. Each of Cassie\u2019s legs has seven DOFs with five of them being actuated: q1, q2, q3, q4, and q7. The corresponding motor torques are u1, u2, u3, u4, and u5. The other two DOFs, q5 and q6, are passive, corresponding to stiff springs. The dynamics of Cassie can then be expressed in the following Euler-Lagrange dynamics: D(q)q\u0308+H(q, q\u0307) = Bu+ JT s (q)\u03c4s + JT c (q)Fc, (3) where q is the generalized coordinate vector as defined in (1), D(q) is the mass matrix, H(q, q\u0307) contains the centripetal, Coriolis, and gravitation terms, B is the motor torque matrix, u is the motor torque vector of dimension 10 corresponding to the actuators on the two legs, Js(q) is the Jacobian for the spring torques, \u03c4s is the spring torque vector, Jc(q) is the Jacobian for the ground contact forces, and Fc is the ground contact force vector",
            " The \u201cfeedback\u201d with \u03b8 and \u03b8\u0307 terms in this equation account for the Hovershoe\u2019s internal stabilization controller that drives \u03b8 to zero when u\u03b8 is zero. The yaw dynamics of the Hovershoe is described by (5), where J\u03c8 is the moment of inertia of the Hovershoe about the z-axis, c3 is the coefficient for the ground contact damping term, and u\u03c8 is the input torque from the rider\u2019s toe about the z-axis. Here, we only have a \u03c8\u0307 term since there is no internal Hovershoe controller driving \u03c8 to zero when u\u03c8 is zero. Figure 3b shows u\u03b8 and u\u03c8 on the Hovershoe. The global translational dynamics of the Hovershoe are captured by (6)-(7), where v is the speed and m the mass of the Hovershoe. Finally, the acceleration dynamics of the Hovershoe is captured by (8) as a function of Hovershoe pitch angle. Note that we did not conduct system identification to determine the parameters of the model. Since computing the tilt angle and applied torque on an accelerating Hovershoe was hard, we instead focused on capturing the structure of the Hovershoe model for use in our control design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001584_amr.1018.15-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001584_amr.1018.15-Figure2-1.png",
        "caption": "Figure 2: Measurement poses for the determination of fastening influences.",
        "texts": [
            " All points are reached from the same direction to eliminate additional error influences. Experiment 2: Isolated consideration of each axis. The selected axis is rotated stepwise while for each step the position of the TCP is measured. Having reached the maximum angle of rotation, the procedure is repeated in the different direction of rotation. The condition of a constant oriented tool is not possible for this experiment. Experiment 3: Influence of robot fastening. To show the influence of the robot fastening, the stretched arm of the robot is moved in different poses (Figure 2). The laser tracker reflector is mounted on the base of the robot to detect deviations due to tilting. The distance of the reflector to the first axis can be estimated to 0.30 m. As a consequence of the reflector position, the effect on the positioning accuracy of the TCP has to be determined by a leverage factor given by the length of the stretched arm. Experiment 4: Dependency on temperature. The influence of temperature on the accuracy is a complex process due to the fact that different mechanisms act on the temperature distribution of the robot elements",
            " Exemplary, the results for the first axis of IR 1 and IR 2 are shown in Figure 5 for normalised working areas. Remarkably, the quotient of absolute accuracy and repeatability for the first axis is in the order of magnitude of 100. Furthermore, the absolute accuracy of the IR 1 robot shows an asymmetric behaviour in the working area (Figure 5 a). For the investigation of the robot fastening, the laser tracker reflector is mounted at the base of the robot. The stretched robot arm generates a torque on the base depending of the robot pose (see Figure 2). Hence, an elastic fastening results in a position shift of the reflector in z-direction. These position shifts are given in Table 2 for the considered robots. The influence on the accuracy of the TCP can be estimated by a multiplication of the values given in Table 2 with the leverage factor taking into account the distance of the reflector to the centre of the robot (first axis). Depending on the robot this effect can lead to positioning errors of the TCP in the order of magnitude of 1 mm. Notably the tilting could not be observed for the IR 2 and only in the range of measurement uncertainty for the IR 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003488_icra40945.2020.9196838-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003488_icra40945.2020.9196838-Figure3-1.png",
        "caption": "Fig. 3: Downwards projection algorithm",
        "texts": [
            " (18) The belly vector closest to the downward direction is the projection of b\u0302 normal to the plane created by w\u0302 and f\u0302, Proj(b\u0302) = f\u0302\u00d7 w\u0302. (19) The three vectors f\u0302, w\u0302, and Proj(b\u0302) are orthonormal and obey the right-hand rule. Aligning the aircraft\u2019s three body axes with these vectors ensures the aircraft will point its thrust to achieve the position control objective, while pointing its belly as downwards as possible. This second condition has as consequence that the vector associated with the direction of the wing, w\u0302 is always contained in the North-East plane. This situation is shown in Fig. 3. Finally, it is straightforward to show that the reference attitude can be expressed as Cri = [\u0302 f w\u0302 Proj(b\u0302) ]T . (20) The projection algorithm encounters a singularity when f\u0302 and b\u0302 become parallel. This will create a problem in (18), since the norm tends to zero. This is a geometric singularity and obeys a clear physical impediment: the aircraft cannot point the nose upwards (or downwards) while pointing the belly downwards. This can be avoided by relaxing the projection condition when the aircraft is sufficiently pitched upwards"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure6-1.png",
        "caption": "Fig. 6 Flexible Cushioning Element",
        "texts": [
            " The rope with 2.5 mm in diameter is driven by a ball-screw pair. The outer shell of the arm is attached with a two-dimensional code to measure the posture of the manipulator by visual fusion method. Due to the redundancy of the rope compared to the DOFs, inconsistent rope lengths caused by control errors will cause the mechanism getting stuck. Therefore, the flexible cushioning element is designed and installed between the ropes and ball-screw pairs to tolerate the inconsistent rope lengths, as shown in Fig. 6. To meet both the requirements of tolerance and accuracy, the stiffness of the flexible cushioning element has been carefully designed and verified. Due to numerous motion joints, To describe and analyze the posture of the snake manipulator prototype, it is necessary to establish an improved kinematic model. The prototype is described using the D-H coordinate system, as shown in Fig. 7. In the figure, Cn is the center of the universal joint of the segment n. Pn(n = 1, 2, \u00b7 \u00b7 \u00b7 5) is the position vector of Cn",
            " The feedback coefficients of PD controller II and PD controller III are shown in Tables 3 and 4, while the control matrix Kp = diag ( kp1kp2kp3kp4kp5kp6 ) and Kd = diag ( kd1kd2kd3kd4kd5kd6 ) are set as diagonal matrix and the feedback coefficients are (Table 5): According to the co-simulation model and the control framework, the simulation errors of the end position and end orientation are shown in Figs. 33 and 34. Although the position error is larger in the first five seconds, up to \u00b11.3mm and \u00b12.5\u25e6 due to the initial state error and the rapidly changing rope velocity, the subsequent error is smaller. Although the flexibility of the virtual spring leads to the stable simulation errors, the errors are small, which can verify the velocity mapping between ropes and end posture. The end drawing circle process is tested using the prototype shown in Fig. 6, and the desired action can be completed, as shown in Fig. 39. In this paper, a red waterbased pen, which is fixed at the end pose of the manipulator, traces the end motion trajectory. The experimental results show that the radius of the end circular path is 0.117m, which is 3.5% different from the expected value 0.113m. The end position error decreases from about 30 mm at the beginning to 2mm in the end in the Fig. 35 and the orientation on error decreases from about 4\u25e6 at the beginning to 1\u25e6 in the end in the Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000948_amm.282.274-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000948_amm.282.274-Figure1-1.png",
        "caption": "Fig. 1 Two rotary motion units (RJ1, RJ2) joined by an arm",
        "texts": [],
        "surrounding_texts": [
            "Introduction\nDIRECT KINEMATICS\nDenavit-Hartenberg\u00b4s principle of coordinate systems location \u2013 location of coordinate systems into segments is arbitrary \u2013 transformation matrix is being created of goniometric relations. Two rotary motion units (RJ1, RJ2) joined by an arm ai, having a general orientation in space with fixed coordinate systems LCSi-1 a LCS\u00ed.\nTransformation relation \u2013 dummy movements to unify both coordinate systems: - rotation of axis Xi-1 at an angle \u03c5, axes Xi-1 and Xi are parallel - displacement of axis Xi-1 in the direction of axis Zi-1 at a distance di, axes are the same - displacement of the beginning LCSi-1 along axis Xi at a distance ai, beginnings LCSi-1 and LCSi are the same - rotation of axis Zi-1 around axis Xi at an angle \u03b1 on axis Zi, LCSi-1 and LCSi are the same - unification by four movements \u2013 rotation, translation, translation, rotation - axes Zi-1 and Zi are axes of rotation RJ1 and RJ2 - axis Xi is a normal to axes Zi-1 and Zi Transformation relation between LCSi-1 and LCSi \u2013 four transformation movements \u2013 transformation matrices in homogeneous form:\nAll rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.216.129.208, University of Auckland, Auckland, New Zealand-21/06/15,23:55:08)",
            "(1)\n(2)\nConsequential transformation matrix:\n(3)\n(4)\nDenavit-Hartenberg parameters: \u03d1 i \u2013 angle between Xi-1 and Xi during rotation around Zi-1 di \u2013 shortest distance (normal line)between axes Xi-1 and Xi, positive direction in direction Zi-1 ai \u2013shortest distance (normal line) between axes Zi-1 and Zi, positive direction in direction Xi \u03b1i \u2013 angle between Zi-1 and Zi during rotation around Xi Rotary unit \u2013 transformation matrix contains only one variable, rotation of unit \u03d1 i, other parameters are constant and characterize other elements dimensions and rotation of motion pairs. Sliding unit \u2013 transformation matrix contains only one variable, shift di, other parameters are constant. Homogeneous transformation matrix \u2013 between two coordinate systems of elements (rotary, translational join) in general form:\n(5)\nR \u2013 sub-matrix of mutual rotation of coordinate systems, p \u2013 displacement vector for beginnings of coordinate systems. Total translational matrix, between basic coordinate system and last n coordinate system, occurs after multiplication of individual translational matrices in order of coordinate systems:\n(6)\nDirect kinematics \u2013 of kinematic pairs q1, q2, ..., qn known displacements dimensions and rotations of kinematic pairs q1, q2, ..., qn:\n(7)"
        ]
    },
    {
        "image_filename": "designv11_30_0001094_j.sysconle.2014.04.008-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001094_j.sysconle.2014.04.008-Figure1-1.png",
        "caption": "Fig. 1. Illustration of the exclusion zone avoidance problem: SC2 must manoeuvre from q0 to qf while vI cam2 avoids vI obs2 \u00b71 by at least \u03c6 \u2200t \u2208 [t0, tf ].",
        "texts": [
            " If the guidance command is not safe, the avoidance part computes a new set of quaternion vectors that avoid the exclusion zones and asks the consensus part to compute a new guidance command in the next step. This cycle repeats until consensus is achieved. Unlike the previous approaches in the literature, the guidance command computed by the consensus part satisfies the quaternion dynamics and the norm preserving constraints while guaranteeing the average consensus (see Section 4.1); and the control input generated by the avoidance part is decentralized and also valid in the multiple coordinate frame setting (see Section 4.2). To illustrate the avoidance part, consider Fig. 1. Denote spacecraft i by SC i and the unit vector of the camera bore-sight by vI cami (t) in F I SCi corresponding to the SC i\u2019s attitude qi(t). Let vI obsi (t) be the unit vector corresponding to the attitude quaternion to the object to be avoided (e.g. bright object or thruster axes). As shown in Fig. 1, we want the time evolution of the camera vector of SC2 from vI cam2 (t0), i.e. q20, to vI cam2 (tf ), i.e. q2f , to avoid vI obs1 (t) at all times, while maintaining a minimum angular separation of \u03c6. In the figure, the vector vI obs1 (t) originating from SC1 is equivalent to vI obs2.1 (t) in the frameof SC2. Thus, the requirement is that the angle \u03b8 between vI cam2 (t) and vI obs2.1 (t) satisfies \u03b8(t) \u2265 \u03c6 \u2200t \u2208 [t0 tf ], or vI cam2 (t)TvI obs2.1(t) \u2264 cos\u03c6 \u2200t \u2208 [t0 tf ]. (1) Note that both vI cam2 (t) and vI obs2",
            " In order to address such practical issues, we develop a mechanism to calculate vI obsi\u00b7j (defined in F I SCi) corresponding to vI obsj (defined in F I SCj). 6 This mechanism is essentially to determine the intersection point of vI obsj (t) with the sphere of radius r , centred at SC i. If indeed there is such an intersection, the intersection point defines vI obsi\u00b7j which can be used to define an attitude constraint represented as (15) and so avoided by SC i. To illustrate, suppose SC1 and SC2 are in their own body coordinate frames relative to Earth, as shown in Fig. 1. A thruster attached to SC1 body frame is at vI obs1 . The circles around SC1 and SC2 are the spheres centred on the origin of their rotation frames. If both spacecraft are close enough, then the vector vI obs1 may pierce the sphere of SC2 at a point, as is shown. This point defines vI obs2\u00b71 in the frame of SC2, and it is desired that as SC2 changes its attitude from q0 to qf , vI cam2 must avoid the exclusion cone created around vI obs2\u00b71 \u2200t \u2208 [t0, tf ]. Note that the above strategy decentralizes the problem in question"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002820_j.jcsr.2020.105959-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002820_j.jcsr.2020.105959-Figure6-1.png",
        "caption": "Fig. 6. Specimen C1 (t2 = 6 mm).",
        "texts": [],
        "surrounding_texts": [
            "The specific geometrical parameters of the test specimens are presented in Table 1 and Fig. 2. The thickness of the side plate is one of themost significant factors that affect the static behavior of the BC joints [18,19]. Hence, three thicknesses of side plate, t2, are considered in the specimens. The bolt diameter in the specimens is 24 mm. The experimental equipment is shown in Fig. 3(a). Three linear variable differential transformers (LVDTs) were equipped for measuring the displacement at points 1\u20133 in order to calculate the joint rotation by formulas (1) and (2). Fig. 3(b) shows the specimen during the test under loading. An apparent deformation appeared on the side plates of the specimen as the load increased. \u03a6 \u00bc P \u03d5ij 3 \u00f01\u00de \u03d5ij \u00bc arctan \u03b4i\u2212\u03b4 j lij \u00f02\u00de \u03b4i and \u03b4j, are the displacement values measured by the LVDTs, and lij is the distance between points i and j on the specimen. The displacement-controlmodewas adopted in this study, as shown in Fig. 4. \u0394y is the displacement value of the end plate as the edge of the side plate yields (Fig. 5), whichwas obtained using FEA in [18]. Each cyclic reverse loading was repeated twice. The specimens were tested until a large plastic deformation occurred or the carrying capacity decreased abruptly. joint."
        ]
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure11-1.png",
        "caption": "Figure 11. Discrete model of the interference area.",
        "texts": [
            " By solving profile equations, the boundary points M1(x,y) and M2(x,y) of the interference area can be obtained, and the equation of a straight line between the two points can be written as y \u00bc ux\u00fe c. By referring to Houpert\u2019s idea of a slicing technique that is used to calculate roller/race contact19, the entire wheel surface is discretized into a number of slices along the x-direction. The width of each slice is x, as shown in Figure 10. The smaller the x is, the more accurate will be the calculation. When geometry interference at point M occurs, the interference area is discretized into many slices. In Figure 11, for the ith slice, the coordinates of its center point Pi can be obtained through a mathematical relationship. A straight line which passes through point Pi and is perpendicular to the line M1M2, can also be acquired. It intersects with the wheel/rail profiles at points N1 and N2, denoted as line N1N2. The coordinates of points N1 and N2 can be obtained by solving simultaneous equations of wheel/rail profiles and the line N1N2. Based on Figure 11, the length of N1N2can be assumed to be the amount of contact deformation of the ith slice. Accurate positions of points N1 and N2 on the wheel/rail profiles can be obtained based on their coordinates. The equivalent radius ratio k(i) of the ith slice is k\u00f0i\u00de \u00bc Rx\u00f0i\u00de=Rz\u00f0i\u00de \u00f03\u00de at RICE UNIV on May 18, 2015pif.sagepub.comDownloaded from at RICE UNIV on May 18, 2015pif.sagepub.comDownloaded from In wheel/rail contact, it can be considered that the normal direction of the contact spot is perpendicular to the line M1M2",
            "3122 0.2117 0.2779 0.1871 at RICE UNIV on May 18, 2015pif.sagepub.comDownloaded from According to equation (12), the density of the sliced load can be introduced as qc\u00f0i\u00de \u00bc dQ\u00f0i\u00de dx \u00bc 4 Plc\u00f0i\u00de 2 blc\u00f0i\u00de \u00bc 4 Plc\u00f0i\u00de 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 8 dQ\u00f0i\u00de dx E Rz\u00f0i\u00de s Rz\u00f0i\u00de \u00f022\u00de The load density of any slice (different i) in an interference region can be calculated using equation (22). In the contact condition shown in Figure 11, assuming that the number of slices in the interference region is n when the wheel profile vertically moves downwards for a distance of y, the load of the interference region can be calculated by Qc \u00bc qc\u00f01\u00de dx\u00fe qc\u00f02\u00de dx\u00fe qc\u00f03\u00de dx \u00fe qc\u00f0n\u00de dx \u00bc Xn i\u00bc1 qc\u00f0i\u00de dx \u00f023\u00de The actual normal load of the contact spot Qw is introduced. Qc is compared to Qw under different y until the following relationship is satisfied Qc Qw 4\" \u00f024\u00de Here the convergence precision \" is very small. If a suitable y is found, the maximum contact stress Pmax under the actual normal load Qw can be obtained by Pmax \u00bc max\u00bdP\u00f01\u00de,P\u00f02\u00de, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002745_s12206-019-1235-8-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002745_s12206-019-1235-8-Figure1-1.png",
        "caption": "Fig. 1. A simplified figure of 3D hybrid manipulator.",
        "texts": [
            " This paper is structured as follows: Firstly, the impedance controller design is presented in which the concept of the virtual foundation is introduced. A control scheme built up for removal of amnesia, which is the accumulated position error during force control, is developed. Ultimately, this impedance control scheme developed is implemented to another application of an inverted mirror L-shaped path on the surface of a submissive object. The schematic diagram of 3D hybrid manipulator is shown in Fig. 1 where two 3D parallel manipulators are connected serially with each other. The non-minimal dynamics of parallel manipulator can be computed using Euler-Lagrange equations [24]. The base is fixed to the ground and it is attached to the moving platform through six legs at points numbered from 1 to 6 in Fig. 1. These six actuators are considered as prismatic pairs. The moving platform of the lower manipulator becomes base for the upper manipulator. This is a six DOF manipulator in which all the joints are taken as pin joints. The centre of gravity G is considered to be in the centre of top moving platform with bottom base fixed. The upper and middle platforms are attached to each other with six legs (actuators) which are considered as prismatic pairs. Similar attachment of middle and lower platforms is required. The lengths of all six legs can be measured and the forces in these legs can be controlled. The leg lengths are calculated by integrating the rate of change of leg lengths that are calibrated using flow sensors in the bond graph representation. The location and direction of centre of gravity of platform are calculated by the values of these leg lengths. xyz coordinate system is connected at the base, middle platform and upper platform as shown in Fig. 1. The directions of x, y and z are taken as same for all coordinate systems. x direction is taken as parallel to the plane of base and ydirection is taken as perpendicular to the plane of base, as displayed in Fig. 1. The angles \u03b8x, \u03b8y and \u03b8z are the Euler angles. The approach of combining a virtual foundation with overwhelming controller has been followed in this paper for impedance control of a three dimensional hybrid manipulator. A compensation gain is used to adjust the impedance at the boundary of robotic arm and environment. The response has some trajectory tracking error from command during the interaction period. This error is known as amnesia and is removed when the manipulator and object are not interacting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000920_gt2013-94087-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000920_gt2013-94087-Figure1-1.png",
        "caption": "Figure 1. Lateral view of advanced bearing test rig (mm).",
        "texts": [
            " This work presents the application of MMFBs in a relatively large rotor configuration representative of a real compressor application, with bearing location modifications as recommended in Ref. [9]. In these experiments, bearing unit loads (static) are 38.8 kPa (5.64 psi) with relatively low surface speeds (42.5 m/s (139 ft/s) at maximum running speed of 9,000 rpm. Care is taken to use best practices as described in the open literature in the manufacturing of the bearings, including the use of coatings and provisions for thermal management. Figure 1 shows the test rig assembly described in detail in Ref. [11], and which is intended to study advanced supports in turbomachinery at real scale. An electric motor drives an intermediate shaft through a multiplying belt and pulley arrangement (ratio of 1:2.63). The intermediate shaft connects to the main testing rotor through a gear-type coupling. The rig allows for different rotor configurations and bearing spans. The current experimental rotor is a 5-stage, 57 kg (without coupling and magnetic bearing sleeves) centrifugal compressor rotor with a balance drum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003004_pesgre45664.2020.9070363-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003004_pesgre45664.2020.9070363-Figure1-1.png",
        "caption": "Fig. 1. (b) The closed loop drive system built in the Simplorer tool showing VSI, HCC and the imported (co-simulated) FE model of BLDC motor.",
        "texts": [
            " Section-III shall illustrate on the modeling of BLDC motor under SITF, demagnetization fault and simultaneous occurrence of both the faults together. Section-IV shall have the detailed discussions on the outcomes obtained through the performance of the machine under the fault conditions. II. FE MODELING OF BLDC MOTOR USING CO-SIMULATION An 8-pole,12-slots inner rotor BLDC motor of 850 W power rating is used for this study, with other parameters as listed in Tables I. The FE model of a BLDC motor is developed using these parameters. The meshed model is shown in Fig. 1. (a) which is simulated for a reference speed of N= 2650 RPM for a frequency, f=50Hz and simulation time t=0.12s. The analysis step size=0.001s with the magnetic mesh approximation = 2mm is given for the machine operation. The electric-drive system comprising of three phase Voltage Source Inverter (VSI) and Hysteresis Current Controller (HCC), is modeled in Simplorer. The FE model of the machine is co-simulated with the Simplorer built drive system operated in a closed loop as given in Fig. 1. (b). The co-simulation is performed under the healthy machine conditions and the characteristics machine quantities are observed. The machine at no load delivers the corresponding B Ministry of Electronics and Information Technology (MeitY), Govt. of India. 978-1-7281-4251-7/20/$31.00 \u00a92020 IEEE 1 phase currents. The no load current of 2.7 A is recorded for a back-EMF of 25V as shown in Fig. 2. (a)-(b). The electromagnetic torque for the corresponding no-load current is 0.6 N-m and the speed under healthy conditions is steady as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003141_j.ijsolstr.2020.06.016-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003141_j.ijsolstr.2020.06.016-Figure1-1.png",
        "caption": "Fig. 1. (a) Vertex with n creases before folding; h1 \u00fe h2 \u00fe \u00fe hn \u00bc 2p. AOB is a straight line perpendicular to crease i with both A and B on the origami and \\AOB \u00bc p. (b) AOB is kinked after folding. \\AOB becomes p qi , where qi is the fold angle of crease i. The triplet, Li;Mi and Ni , defines a local Cartesian coordinate system for the sector facet i; i\u00fe 1\u00f0 \u00de: Li is the unit vector along crease i;Ni is the unit normal of the sector facet i; i\u00fe 1\u00f0 \u00de and Mi \u00bc Ni Li where \u2018\u2018 \u201d is the cross product. The indexes (i-1) and (i + 1) are to be interpreted cyclically.",
        "texts": [
            " The loop closure constraint has to be fulfilled at every interior vertex for all the rigid facets to be compatible with each other during the folding process (Belcastro et al., 2002; Lang and Twists, 2017). Wu and You investigated the folding of rigid origami based on the rotating vector model, which describes the loop closure constraint using quaternions (Wu et al., 2010). For a given crease pattern, the three-dimensional (3D) folded form of the rigid origami can be uniquely determined by the fold angles, which are the supplementary angles of the dihedral angles between two adjacent facets (Fig. 1). Kinematic folding of rigid origami can be described by parametric equations for origami tessellations comprised of identical unit cells or for origami with a single vertex. Wei et al. analytically calculated the Poisson\u2019s ratio and stiffness of the Miura-ori tessellation, based on the geometry of the unit cell (Wei et al., 2013). Assuming symmetry between the fold angles, Hanna et al. investigated the kinematics of a degree-8 Waterbomb base, consisting of a single vertex with 8 symmetric creases (Hanna et al",
            " In addition, by considering the rotational springs at the creases, an algorithm based on the Lagrange multiplier method is presented to search for the equilibrium configuration of the rigid origami. It is shown that the projection method by Tachi (2009) agrees with a special case of the current physical model. The algorithms are then verified by several examples. The rigid origami model of Tachi (2009, 2012) is reviewed here for completeness. The origami is commonly designated by the crease pattern, consisting of vertexes and creases. Vertexes are points on the origami paper and each crease is a line joining two neighboring vertexes along which the origami paper is folded. Fig. 1(a) shows an isometric view of unfolded origami paper of a typical vertex, P, with n creases joining it. The creases are numbered from 1 to n anticlockwise. Creases i and i\u00fe 1 define the sector angle hi. It is trivial that the sum of the n sector angles is 2p. Before folding, AOB is a straight line perpendicular to crease i on the origami paper, i.e., \\AOB is p as shown in the figure. After folding, \\AOB would be denoted as p qi, in which qi is the fold angle of crease i. For the rigid facets to be compatible with each other, the loop closure constraint around vertex P is F qP\u00f0 \u00de \u00bc v1;2v2;3 ",
            " 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. This work is supported by the Key Research and Development Program of Anhui Province (1804a09020001). This appendix first derives the loop closure constraint; the derivative of the constraint matrix with respect to fold angles is then proven to be antisymmetric for compatible fold angles. In Fig. 1(b), the triplet, Li;Mi and Ni, defines a local Cartesian coordinate system for the sector facet i-(i+1) with Li being the unit vector along crease i;Ni being the unit normal of the sector facet i- i\u00fe 1\u00f0 \u00de and Mi \u00bc Ni Li where \u2018\u2018 \u201d is the cross product. It is clear that Li; Mi; Ni\u00bd \u00bc Li 1; Mi 1; Ni 1\u00bd vi 1;i \u00f0A:1\u00de where vi 1;i given in Eq. (2) for i \u00bc 1;2; . . . ;n and (i-1) is to be interpreted cyclically. Starting from the sector facet 1-2 and looping around the vertex anticlockwise, recursive usage of the transformation in Eq",
            " To ensure an ordered calculation, a spanning tree can be constructed on the crease pattern with its root at the facet whose position is specified (Lang and Twists, 2017) (Fig. B.12). For instance, we consider the calculation of the coordinates of vertex-1 whilst only the relevant vertexes (black numbers) and creases (blue numbers) are indicated in Fig. B.12. First, the facet 7-6-1-2 is to be rotated with respect to facet 8-7-2-3. The axis of rotation is the vector V2V7 ! . Owing to the sign conventions of the fold angles in Fig. 1(a), the order of the vertexes for the axis should be in the clockwise direction regarding the facet 8-7-2-3 which is nearer to the root facet 12-11-9-10. Let qi be the fold angle at crease-i; ai; bif g be the pair of ordered vertex numbers for crease-i such that VaiVbi ! is in the clockwise direction of the facet that is near the stationary facet; and Xi \u00bc xi; yi;0f g be the coordinates of the crease pattern vertex-i. Then x1 \u00bc Ra qi; ei\u00f0 \u00de x1 Xai\u00f0 \u00de \u00fe Xai with ei \u00bc Xbi Xai jXbi Xaij \u00f0B:1\u00de where Ra qi; ei\u00f0 \u00de is the rotation matrix which rotate a vector by angle qi about the unit vector ei (see Lang and Twists, 2017 for the relevant expressions) and xi \u00bc x; y; zf g donates the coordinates of vertex-i in 3D space"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002979_tii.2020.2986805-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002979_tii.2020.2986805-Figure5-1.png",
        "caption": "Fig. 5. The scheme for insertion ratio. (a) Vertical insertion. (b) Inclined insertion.",
        "texts": [
            " Since the object and its manipulator are connected by a spring, the manipulator\u2019s insertion does not naturally equal the object\u2019s insertion. A parameter is defined to represent this difference \u03c3v,t = \u2206vo,t \u2206vg,t = { 1+ Lv,t\u2212Lv,t\u22121 \u2212\u2206vi,t \u00b7iz , sv = 1 1\u2212 Lv,t\u2212Lv,t\u22121 \u2206vw,t \u00b7iz , sv = 0 (31) where \u03c3v,t is the insertion/withdrawal ratio between the motion of the upper object and its manipulator in vertical insertion, \u2206vo,t and \u2206vg,t are the insertion/withdrawal depths of the object and its manipulator, as shown in Fig. 5(a). This ratio is an important factor that affects future insertion planning. There are four conditions for different insertion/withdrawal ratio considerations. 1) In the initial insertion process, the spring length has not yet been compressed to an extent, and the object keeps still or moves only a little, i.e., \u03c3v,t < \u03b5\u03c3 exists, where \u03b5\u03c3 is a small parameter. In this sense, the maximum insertion depth is chosen and the manipulator insertion parameter is \u03bev,t = 1. Authorized licensed use limited to: University of Liverpool",
            " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. projection onto the estimated inclination \u03c3i,t = \u2206ao,t \u2206ag,t = { 1+ (Lt\u2212Lt\u22121)\u00b7dt \u2212\u2206ai,t \u00b7dt , si = 1 1\u2212 (Lt\u2212Lt\u22121)\u00b7dt \u2206aw,t \u00b7dt , si = 0 (38) where \u2206ao,t and \u2206ag,t are the insertion depths of the object and the manipulator, independently, for the current time step, and \u03c3i,t represents the insertion/withdrawal ratio between object and manipulator for inclined insertion, as shown in Fig. 5(b). There are also four similar stages for inclined insertion. 1) The first stage is the initial insertion process where the spring is compressed and the object does not insert. The inclination estimation has not yet started and the maximum insertion depth is applied with \u03bei,t = 1. 2) After the object moves, the ratio is \u03c3i,t \u2265 \u03b5\u03c3 , and resorting to (32), the estimated ratio of the next step \u03c3 i,t+1 is computed by replacing \u03c3v,t with \u03c3i,t . Since the inclination needs to be estimated online, the next insertion depth of the object leads to \u2206a\u0303o,t+1 = |1\u2212 kd\u2225dt \u2212dt\u22121\u22252| ( 1\u2212 \u2225Fr,t\u22252\u2212\u2225Fr,t\u22121\u22252 k f ) \u2206ao,t , (39) where \u2206a\u0303o,t+1 is the object\u2019s computed insertion for the next time step before limitation and kd scales the inclination estimation error"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002780_j.optlaseng.2020.106039-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002780_j.optlaseng.2020.106039-Figure5-1.png",
        "caption": "Fig. 5. (a) 3D infrared image (one image of the recording sequence) (b) RGB image with restored wing section (polymer modelling paste) in pink (c) Automatic determined defects/difference in material (red) with a standard PCA analysis (with the background removed). (d) Automatic determined defects/difference in material with the proposed weighted PCA. This weighted method does not include the false positives of the standard PCA method.",
        "texts": [
            " Experimental results The detection methodology is tested on a cast figurine of an angel nd a prototype bicycle part. The measurements of the bicycle part show hat the quality map can be used to avoid to detect false defects. .1. Cast figurine The proposed method is tested on a restored cast figurine of an angel. nitially, the figurine right wing was missing and is sculptured with poly- er modelling clay. Next, the angel was repainted with acrylic paint to ide the restoration. The ground truth polymer clay restoration is high- ighted in pink on the image in Fig. 5 (b). The 3D shape is obtained by canning the figuring with a structured light scanner (Faro Scan-In-A- ox). The heat response of the figurine is recorded, and next the 3D pose is alculated. A mapped heat image on the 3D model is visible in Fig. 5 (a). fterwards, the quality map is calculated, and the heat response is analsed with the PCT analysis (with background removed) ( Fig. 5 (c)) and he proposed PCT analysis using weights ( Fig. 5 (d)). The red parts are parts that correspond with a high chance of diference in material properties (highest 1% values in the second temoral mode). In the standard PCT analysis ( Fig. 5 (c)) parts of the head rm,legs and foot are highlighted incorrectly. The standard analysis also ighlights surface edges. In this case, this is undesirable because these dges do not correspond with a change in material properties. In the mage showing the weighted analysis using the quality map ( Fig. 5 (d)) hese edges are not highlighted. In contrast to the standard method, the roposed analysis highlights the wing correctly without false positives. he restored section has an area of 4705 pixels that is highlighted in he infrared image. The standard method highlights 3814 pixels, from hich 2341 pixels are correctly highlighted (50% of the restored area). he proposed method only highlights 2015 pixels of the true 4705 pixels 42%), which is less than the standard method but has no false positives"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002510_cec.2016.7744141-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002510_cec.2016.7744141-Figure1-1.png",
        "caption": "Fig. 1. Depiction of earth and body-inertial frames [30]",
        "texts": [
            " Section VI presents simulation results with discussion and the paper is concluded in section VII. A quadrotor is an underactuated system with 4 motors which are essentially the system inputs. The model itself is a 6 degrees-of-freedom (DOF) system with attitudes to maintain system stability. These include the 3 translational terms for forward (x), horizontal (y) and vertical (z) movement as well as the yaw (\u03c8), pitch (\u03b8) and roll (\u03c8) rotations around these aforementioned translational axis. These DOF are graphically shown in Fig. 1. As in the case of all flying vehicles, multiple reference frames are established and an Earth and body-fixed frames 2792 2016 IEEE Congress on Evolutionary Computation (CEC) are chosen in this specific case. Modelling of quadrotor systems requires such formulations to ensure that Newton\u2019s laws are applicable and also permit a simple formulation with minimal complexities. The body-fixed frame coincides with the centre of gravity (COG) of the quadrotor and the axis of this frame follow the body in its forward (x), horizontal (y) and vertical directions (z). It also includes the rotational freedom across these respective axis as well, namely (\u03c8, \u03b8, \u03c8). The Earth-fixed frame is essentially the realization of the quadrotor motion relative to a stationary body (Earth). It relates the motion of the quadrotor to the forward (X), horizontal (Y ) and vertical (Z) directions of the observer itself. The concept of the reference frames is graphically depicted in Fig. 1. Transformations between these frames are achieved through transformation matrices R and T , which achieve transformations relating to the translational axis and rotational planes respectively. These are defined as follows [30]: R = \u23a1 \u23a3 c\u03c8c\u03b8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03b8 c\u03c8s\u03b8s\u03c6 + s\u03c8s\u03b8 s\u03c8c\u03b8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03b8 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03b8 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8s\u03c6 \u23a4 \u23a6 (1) T = \u23a1 \u23a3 1 t\u03b8s\u03c6 t\u03b8c\u03c6 0 c\u03c6 \u2212s\u03c6 0 s\u03c6/c\u03b8 s\u03c6/c\u03b8 \u23a4 \u23a6 (2) where s, c and t are the sine, cosine and tangent functions of the respective Euler angles. The translational and rotational co-ordinates in the Earth-fixed frame are [30]: \u0393E = [ x y z ], and \u0398E = [ \u03c6 \u03b8 \u03c8 ]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003214_jmems.2020.3005090-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003214_jmems.2020.3005090-Figure4-1.png",
        "caption": "Fig. 4. CAD design of the device.",
        "texts": [
            " Wire cut EDM, using a 400\u03bcm diameter wire is then used to detach the printed components from the build platform. The components are then rinsed in isopropyl alcohol and dried before further processing. Figure 1(c) shows the released devices. Scanning electron microscopy was used to characterize the as-printed components, Figure 3. The first step towards the transfer of the MEMS component on a substrate is cleaning. A soda lime glass slide substrate was used for the transfer. Before the transfer, the glass surface was cleaned using lab-grade cleaning paper. Figure 4 shows the CAD design of the MEMS switch to be transferred. The green surfaces of the device are the anchor points. Figure 4e and 4f shows the side and cross-sectional view of the device with green anchor points. The fixed beam, as seen, are thicker than the cantilever beam. The green points in Figures 4e and 4f are to be attached to the final substrate leaving the cantilever beam freely suspended. The red areas indicate the sacrificial metal frames which orient the structure during processing and will be removed later for electrical isolation of the switch, much like a sprue in injection molding. Once the device is mounted onto the epoxy and mechanically held by PMMA glue, the removal of the sacrificial metal frame will not lead to the loss of the desired orientation and gaps between electrodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003563_j.ijleo.2020.165776-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003563_j.ijleo.2020.165776-Figure1-1.png",
        "caption": "Fig. 1. Physical model of finite element simulation of the valve seat.",
        "texts": [
            " (5) [21]: {d\u03c3} = [ Dep ] {d\u03b5} \u2212 [Cth][M][\u0394T] (5) where d\u03c3 was the stress increment; d\u03b5 was the strain increment; [Dep] was the elastoplastic stiffness matrix; [Cth] was the thermal stiffness matrix; [M] was the temperature shape function; [\u0394T] was the temperature change amount. During the stress field simulation, a displacement constraint was applied to the bottom of the model to prevent rigid rotation and translation of the elastoplastic model. In order to accurately obtain the residual stress, the temperature load for stress calculation was the temperature from high temperature to cooling to room temperature. F.-Z. Sun et al. Optik 225 (2021) 165776 Fig. 1 shows the physical model of finite element simulation constructed based on the real valve seat of the new type of power. The inner and outer diameters of the valve seat were \u03a632 mm and \u03a648 mm, the corresponding outer diameter of the cone surface was \u03a640 mm, and the model height was 6 mm. The grid size was 0.4 mm, and the grid was obtained through pre-optimization. The movement of the laser heat source model on the valve seat was realized through the secondary development of the APDL language of ANSYS software"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002253_1464419316660930-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002253_1464419316660930-Figure1-1.png",
        "caption": "Figure 1. Scheme of the rolling element \u2013 ring contact.",
        "texts": [
            " Thus, each RE gets loads due to the contact with the rings and the cage, each ring gets loads because of the contact with each RE and the cage, and the cage gets loads from its contact with each RE and each ring. What follows is the way in which these contacts are modelled. Rolling element \u2013 ring contact. The contact force due to the contact between the REs and the rings can be divided into two forces depending the direction in which they are applied. Thus, a force normal to the mating surface in which the contact is produced and a force lying on the contact plane are obtained. The directions of these forces are given by the unitary normal vector n\u0302 and the unitary tangential vector t\u0302 shown in Figure 1. It should be noted that the vectors in Figure 1 represent the unitary vectors for the load supported by the REs. Hence, the loads supported by the rings have the opposite direction. The normal force is calculated as the sum of elastic and dissipative components.53 The elastic force is determined by means of Hertz theory.54 According to that, deformations occur in the elastic range and the dimensions of the contact area are small compared to the radii of curvature of the bodies under load. Note that for the normal force a dry contact is assumed, which nearly approximates EHL conditions for the highly loaded zone (as stated by Rahnejat and Gohar22) but it is not so accurate for lightly loaded zones where clearances emerge",
            " The expression to calculate Kn is the following Kn \u00bc 1 1=Ki\u00f0 \u00de 1=n \u00fe 1=Ko\u00f0 \u00de 1=n n \u00f05\u00de where Ki and Ko are the stiffness of the contacts between the REs and the inner and outer rings, respectively. These two values are functions of the geometrical and material properties of the bodies, and the procedure to calculate them is explained by Harris and Kotzalas.54 Regarding the deformation, it is equal to zero if there is no contact and it is calculated as the distance between the nearest surfaces of both the RE and the raceway of the ring otherwise, as it is shown in Figure 1 for a ball bearing. For the case of roller bearings, the contact is assumed to be finite length and it should be highlighted that the value of the deformation may vary along the roller length when misalignment is produced, taking into account the effect of roller tilting.55,56 Regarding the tangential force, it takes into account the effect of the lubricant and the rolling/sliding motion of the REs. According to the EHL theory, the shear stress in the lubricant film is defined as a function of the lubricant viscosity and the strain rate _ , i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000049_s00170-019-04441-3-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000049_s00170-019-04441-3-Figure4-1.png",
        "caption": "Fig. 4 a CFD powder flow simulation results and b experimental powder flow results",
        "texts": [
            " In addition, a wall boundary condition was set on the surrounding walls of the powder nozzle and protective gas nozzle. A pressure outlet boundary condition was set on the cylindrical zone covering the volume beneath the nozzle [1]. Table 1 summarizes the powder and processing parameters considered in the DED simulations. In accordance with the specification of the experimental nozzle head, the diameter of the powder inlet was set as 3 mm and the diameter of the protective gas flow was set as 10 mm. Figure 4a and b compares the CFD and experimental results for the powder-gas interaction beneath the coaxial nozzle. In both cases, the powder-gas flow spreads freely in the air, and the flows from the nozzles intersect at a point directly beneath the coaxial nozzle. An inspection of the CFD simulation results shows that the powder mass concentration ranges from 0.2 to 0.94 kg/m3 and forms a cylindrical volume along the vertical axis. Moreover, the maximum powder mass concentration occurs at a distance of 15",
            " Figure 14 presents the CFD results for the profile of powder concentration for a fixed powder feeding rate of 18 g/min and four different powder gas flow rate from 5 to 20 L/min. The increasing of the powder-gas flow rate does not bring significant change to the dimensions of the powder concentration profile and the value of the powder concentration. The maximum powder concentration decreases approximately 6.4% when the powder-gas flow rate increases from 5 to 20 L/min. Dimensions of the concentration profiles observed very insignificant change with increasing powder-gas flow rate. In Fig. 4, the average width and height are recorded approximately to be 0.72 mm and 1.37 mm, respectively. To investigate the effect of the powder-gas flow rate on the absorptivity of the system, simulations were carried out for a fixed powder feeding rate of 18 g/min and various powder-gas flow rates in the range of 5~20 L/min. As shown in Fig. 15, the absorptivities of the powder particles and substrate undergo only a minor change as the powder-gas flow rate increases. In other words, the powder particle velocity has only a small effect on the absorptivity of the DED system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002103_j.apm.2016.04.029-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002103_j.apm.2016.04.029-Figure1-1.png",
        "caption": "Fig. 1. Scorbot robot control system.",
        "texts": [
            " The designed parameter-varying observer is not only used to detect faults but also to estimate unmeasured states. Especially, when k i ( t ), i = 1 , . . . , n are time invariant, and = \u03be , the proposed observer will be reduced to the classic Luenberger observer. Thus, the proposed approach is more complicated and more practical than the aforementioned papers. 5. Simulation example To illustrate the validity of the proposed fault detection and accommodate control, the Scorbot robot ER-VII manipulator system (see Fig. 1 ) is simulated, which is described by the following differential equation [36] : M \u0308q + G (q ) + F f (q, \u02d9 q) + F L = nk t i L m di dt + R m i + k b \u0307 q = V, (45) where M = (m 1 + m 2 ) L 2 / 3 , G (q ) = (m 1 + m 2 ) Lg cos q = G N (q ) + G (q ) , F f (q, \u02d9 q) = [ F c + (F s \u2212 F c ) exp (\u2212| \u0307 q/ v s | 2 f )] \u00b7 sgn ( \u0307 q) = F f N + F f , q represents joint position; F f (q, \u02d9 q) \u2208 R represents friction torque; F L \u2208 R represents external disturbance; G ( q ) and F f denote the unknown uncertainties of M, G ( q ) and F f (q, \u02d9 q) , respectively; i represents the motor current The values of the robot manipulator and actuator parameters are borrowed from [36] and represented in Table 1 ",
            " The observer design parameters are chosen as k\u0304 1 = k\u0304 3 = 2 , k\u0304 2 = 10 and = 50 ; the controller design parameters are chosen as c 1 = c 2 = 5 and c 3 = 20 , \u03b3i = 10 , \u03b7i = 1 , \u03c3i = 1 , \u03bci = 0 . 5 , i = 1 , 2 , 3 . The performance functions are set as \u03b7 = 0 . 3 e \u22122 t 1 + 0 . 2 and \u03b4 = \u03b4\u0304 = 0 . 3 , and the threshold-modification parameters are chosen as \u03b5 = \u03b5\u0304 = 0 . 05 . The reference output is taken as y d = sin t . The initial conditions are selected as x (0) = [0 . 2 , 0 . 1] T , and other initial conditions are assumed to be zero. The simulation results with or without FA are shown in Fig. 2 (a)\u2013(d). The fault detection errors and their threshold are shown in Fig. 1 . From the proposed FD scheme, the fault is detected almost immediately at approximately T d = 5 . 35 s as shown in Fig. 2 (a). After the fault detection T d = 5 . 35 s, the proposed adaptive accommodation controller is used, and the design parameters of the observer is adjusted as k\u0304 1 = k\u0304 3 = 5 , k\u0304 2 = 20 , and the performance function are reset as \u03b7\u2032 = 0 . 1 e \u22122(t\u22125 . 35) + 0 . 01 and \u03b4\u2032 = \u03b4\u0304\u2032 = 1 . To show the fault accommodation performance with the proposed adaptive controller, the simulation using the nominal controller that continues to work after the FD is performed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001210_ecce.2014.6953930-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001210_ecce.2014.6953930-Figure9-1.png",
        "caption": "Fig. 9. Structure of the test drivetrain consisting of a two-stage gearbox connected to a PMSG with characteristic vibration frequencies.",
        "texts": [
            " It should be pointed out that the values of and are limited into certain ranges, to assure the generalization capability of the SVM. In this work, the ranges are set to be 0.1, 100 and 0.1, 1000 . IV. EXPERIMENTAL VALIDATION The proposed hybrid classifier is validated for diagnosis of multiple types of faults in a test drivetrain gearbox. Fig. 8 shows the experimental system setup, which consists of a 300-W PMSM driven by a variable-speed induction motor (IM) through two back-to-back connected gearboxes. These two gearboxes are two-stage gearboxes with 3 shafts and 4 gears. Their internal structure is shown in Fig. 9. The test gears are mounted at the input shaft of the test gearbox and pretreated by artificially generating various faults which were commonly observed in industrial systems, including one-tooth missing, two-tooth missing, and a gear crack (Fig. 10). One phase stator current of the PMSM is recorded with a sampling rate of 10 kHz for 210 seconds, during which the rotating speed of the PMSM is varied randomly in the range of 297 to 891 rpm. A total number of 1326 data records are collected, including 340 healthy cases, 318 one-toothmissing cases, 293 two-tooth-missing cases, and 375 gear crack cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001179_s12541-014-0399-5-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001179_s12541-014-0399-5-Figure2-1.png",
        "caption": "Fig. 2 Design of laser head",
        "texts": [],
        "surrounding_texts": [
            "\u00a9 KSPE and Springer 2014\nStudy on Fiber Laser Welding Conditions for the\nFabrication of a Nuclear Fuel Rod\nJintae Hong1,#, Chang-Young Joung1, Ka-Hye Kim1, Sung-Ho Heo1, and Hyun-Gil Kim2\n1 Neutron Utilization Technology Division, Korea Atomic Energy Research Institute, Daedeok-daero 989-111, Yuseong-gu, Daejeon, Korea, 305-353 2 LWR Fuel Technology Division, Korea Atomic Energy Research Institute, Daedeok-daero 989-111, Yuseong-gu, Daejeon, Korea, 305-353\n# Corresponding Author / E-mail: jthong@kaeri.re.kr, TEL: +82-42-868-4420, FAX: +82-42-868-8364\nKEYWORDS: Nuclear fuel rod, Fiber laser welding, Shield gas, Nuclear fuel irradiation test\nTo verify the performance of nuclear fuel, a test rig needs to be fabricated to place the nuclear fuel and carry out irradiation tests. Welding is one of the most frequently used processes in fabricating a test rig, and a fiber laser welding system that enables assembling the components with various volumes has been developed in this study. A case study was then carried out for a nuclear fuel rod made of Zircaloy-4 by changing the process variables such as the focal length, frequency of laser pulse, peak laser power, and shield gas. From the case studies, a suitable welding condition with which a fuel rod can be completely welded regardless of the shield gas (He or Ar) should be achieved. By inspection of SEM and an immersion test, it was verified that Ar is useful as a shield gas in fabricating a nuclear fuel rod with a fiber laser welding system.\nManuscript received: December 30, 2013 / Revised: January 28, 2014 / Accepted: February 5, 2014\nTo verify the performance of a newly developed nuclear fuel, an irradiation test needs to be carried out in the loop of a research reactor, which simulates the condition of a reactor core of a nuclear power plant.1 In addition, a test rig in which nuclear fuel is installed needs to be fabricated to carry out an irradiation test. In particular, the welding machine is one of the most important types of equipment for assembled sensors and components consisting of a test rig.\nPrevious studies to fabricate a test rig applied TIG welding, magnetic force electrical resistance welding and ND:YAG laser beam welding in fabricating a test fuel rod, and some studies applied EB welding to assemble small components such as sensors and their fixtures.2-5 However, because TIG welding and resistance welding techniques have\nsome disadvantages such as a high oxidation problem, wide heat affected zone (HAZ), difficulty in welding dissimilar metals, and poor dimensional precision, they are not suitable techniques in assembling components with a complicated geometry. In addition, although the Nd:YAG laser beam welding technique shows a high dimensional precision, it has certain disadvantages such as low efficiency of laser oscillation (efficiency 4~5%), high heat generation while operating a welding machine, and a short component lifetime. The EB welding technique is mainly used in assembling precise components. However, because it shows good performance only in a vacuum chamber, it is not suitable for assembling large components.\nFiber laser welding techniques have recently been developed and broadly utilized in heavy industries, electronics, semi-conductor industries, and medical science owing to several advantages such as high efficiency of laser oscillation, little heat generation during operation, and a small volume of equipment.6\nIn this study, an automatic fiber laser welding system was developed to assemble an irradiation test rig that consists of components with various shapes and various sizes including dissimilar metals. In particular, previous studies that apply the TIG welding technique and Nd:YAG laser welding technique in fabricating a nuclear fuel rod have applied He gas as a shield gas because it is inert gas with a high heat transfer rate owing to its high mobility. However, He is expensive and\nNOMENCLATURE\nLoffset = Offset distance from the focal length Ppeak = Peak laser power fpulse = Frequency of pulse\nDOI: 10.1007/s12541-014-0399-5",
            "some He can be trapped in the bead during the welding process because the volume of a He molecule is too small, which causes a growth of corrosion during the irradiation test. Therefore, Ar gas (three- or fourtimes cheaper than He) is considered as the second source of the shield gas, and its welding characteristics will be inspected in this study.\nAs shown in Fig. 1, a fiber laser welding system was developed, which consists of five parts such as a fiber laser source, a laser head, a rotational index chuck, a three-axis linear motion guide, and a control PC for the CNC controller.\nQCW 150-1500-AC made by IPG Co. Ltd. is used as a fiber laser source, which is an air-cooled type, and the wave length of the laser is 1070 nm. Both pulsed laser mode (QCW) with 1500 W of peak power and continuous laser mode (CW) with 250 W of peak power are available.\nThe laser head is equipped on the Z-axis LM guide. It was designed with an anti-reflection structure to minimize damage to the optical fiber owing to the reflected beam generated during the welding process, and was also designed to supply shield gas through the nozzle installed at the outlet of the laser beam. The focus lens was designed to have a 180\nmm focal length and \u2205300 \u00b5m beam diameter (measured beam diameter: \u2205276 \u00b5m). In addition, a CCD camera is equipped on the axis of the beam path to enable the operator to find the exact welding position and monitor the generation of welding beads in a timely manner.\nA rotational index check is installed on the workbench, and designed to rotate the workpiece through CNC control. In addition, the chuck can fix tubes with a diameter of \u22051 to \u220568 mm.\nThe three-axis LM guide is equipped on the workbench, and its transfer is controlled by the CNC program with a precision of the Xaxis (900 mm \u00b1 0.02 mm), Y-axis (300 mm \u00b1 0.02 mm), and Z-axis (200 mm \u00b1 0.02 mm). Three axes can be controlled simultaneously, and circular components and inclined components can be welded on a single path. In particular, because both ends of the base frame are designed to be opened, and the rotational index chuck is hollow type, the system can fix a 5-meter-long component and carry out welding accurately. The control PC controls the welding head, shield gas, and three-axis LM guide according to the designated welding path and laser power. The welding position, path, and beads can be checked through the separated monitor whose image is transferred from the CCD camera.\nThe developed fiber welding system is applied to test the welding performance of the nuclear fuel clad and the end cap made of Zircaloy4 (Table 1). A nuclear fuel rod is fabricated by putting nuclear fuel pellets, alumina pellets, and plenum into the fuel clad, and sealing the fuel clad with two end caps (Fig. 3). Their welding will be carried out in a circumferential direction by rotating the fuel clad with 0.087 rad/ sec of angular velocity.\nPreliminary experiments were carried out using CW mode and QCW mode at the focal length (180 mm) between focal lens and the workpiece. As shown in Fig. 4, sparks frequently fly up and the bead is not generated smoothly because some of the molten metal is burnt out or flows out of the bead. Because the peak power of a laser beam is configured at the focal length, the molten metal generated by the previous beam spot absorbs the beam energy at the overlapped area, and some of the molten metal is burnt out. Therefore, the offset distance from the focal length is considered to relieve the heat concentration on the surface area of the workpiece.\nSn Fe Cr O Si Zr\n1.24 0.21 0.11 0.13 0.008 Bal.",
            "In this study, three cases of experiments were considered to find out the optimal process variables of the fiber laser welding process as shown in Table 2. The developed fiber laser welding system has four process variables such as the offset distance from the focal length (Loffset), peak laser power (Ppeak), and frequency of pulse (fpulse) including the shield gas. The first case is checking the quality of the welded surface according to the offset distance from the focal length (Loffset). The second case is checking the quality of the bead according to the frequency of the laser pulse. The last case is checking the penetration depth according to the peak laser power. The above three kinds of experiments were carried out using He and Ar as a shield gas, respectively. According to the preliminary experiments, the flow rate and pressure of the shield gas are fixed with 20 liter/min. and 0.15 MPa, respectively.\nThe results of the fiber laser welding experiments by changing the offset distance from the focal length (Loffset) are shown in Fig. 5. At this time, the frequency of the pulse and the peak laser power are fixed with 10 Hz, and 750 W, respectively.\nWhen He is applied as a shield gas, welded beads at 3 mm or more of Loffset have a uniform surface quality without oxidation film. In addition, when Ar is applied as a shield gas, welded beads at 5 mm or more of Loffset have a uniform surface quality with a little oxidation film. The width of bead when Ar is used as a shield gas is on averagely 1.66% wider than that of a bead when He is used. The above difference is based on the difference of heat transfer rates between two shield gases, in which He molecules are lighter than Ar molecules, and they have higher mobility than Ar molecules. It is obvious that the power density at the welding bead becomes higher as the offset distance from the focal length is shorter. According the above experimental results,\nthe shortest Loffset, which shows a uniform bead in both cases which apply He and Ar as a shield gas, is 5 mm.\nAccording to the result in the previous paragraph, Loffset and Ppeak are fixed at 5 mm and 750 W in this paragraph, respectively. The results of fiber laser welding experiments by changing the frequency of laser pulse (fpulse) are shown in Fig. 6. Both cases applying He and Ar as a shield gas show fine and uniform beads at 8 Hz, 10 Hz, and 12 Hz. Beads generated under 8 Hz of fpulse do not fully overlap the neighbor beads, and it cannot completely seal out the clad with a uniform penetration depth. Beads generated at over 12 Hz of fpulse overflow into the neighbor beads because the next pulse is induced on the workpiece before the molten metal generated by the previous pulse generates a bead.\nAccording to the results in the previous paragraphs, Loffset and fpulse are fixed at 5 mm and 10 Hz, respectively. The results of fiber laser welding experiments by changing the peak power (Ppeak) are shown if Fig. 7. In this experiment, a section inspection was carried out to check"
        ]
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.25-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.25-1.png",
        "caption": "Figure 2.25 Flexible electrodes (A) with Mylar film (DuPont, Mylars) and (B) of conductive rubber.",
        "texts": [
            " A flexible electrode is fabricated depositing a metallic film, for example, Ag AgCl, on a flexible substrate (polymeric); usually the substrate side receives a layer of conductive adhesive to fix the electrode to the skin, and the metallic face, which is connected to a lead wire, receives a layer of insulating material. The electrode must be thin, flexible, and easy to adjust to the contours of the body and stay at the place for long periods. They are used primarily for monitoring children and premature newborns because they are light, thin, and adapt to body contours, avoiding motion artifact and ulceration in the contact region, which may occur in the case of rigid electrodes applied with electrolyte and adhesive tapes. The electrode shown in Figure 2.25A was developed in the 70s for use in newborns. The substrate is made from a very thin Mylars film (13 \u03bcm thickness) coated by an Ag layer (1 \u03bcm thickness); then a thin layer of AgCl is deposited electrolytically giving the electrode an appearance of mesh. The flexible electrode fabricated with Mylars substrate, a polyester film invented in the early 1950s (DuPont, Mylars), is also transparent to X-rays, and this means that the patient may be subjected to X-ray examination without removing the electrode from the skin, which minimizes skin irritation, mainly in babies. Other polymeric materials used as flexible electrode substrate are PDMS (polydimethylsiloxane), Parylene, and Polyimide (Baek, An, Choi, Park, & Lee, 2008; Ochoa, Wei, Wolley, Otto, & Ziaie, 2013; SCS Specialty Coating Systems). Figure 2.25B shows another type of flexible electrode, the conductive rubber electrode, made from silicone rubber with embedded carbon particles. This electrode is often used as stimulation electrode, for instance, in the application of transcutaneous electrical nervous stimulation or TENS. The fixation of flexible electrode is made with a thin, also flexible layer, the same size of the contact surface of the electrode, of a gelatinous material, called electrolytic hydrogel. It is adhesive (minimizes the motion artifact) and conductive and its electrical impedance is higher than that of the conventional electrolytic gel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001323_1.4024555-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001323_1.4024555-Figure2-1.png",
        "caption": "Fig. 2 Typical case. From left to right: elastic deformation, contact pressure, and hydrodynamic pressure",
        "texts": [
            " If the pressure of a quadratic element node is found negative, the element is divided in four linear elements, the central node being obtained with the element interpolation function. From experience, most of cases are successfully treated with a 4 128 TEHD grid; four elements for half width and 128 elements on the circumference. All of the necessary technical details, dealing with Finite Element TEHD modeling, are given by Bonneau et al. [18]. The typical computing time is about 4 h on a common desk computer, which is quite long for an ambitious parametric study. Figure 2 presents a typical severe case: near the maximum normal load (90 kN, 360 deg of crankshaft), at 2000 rpm, the 041704-2 / Vol. 135, OCTOBER 2013 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 11/28/2013 Terms of Use: http://asme.org/terms hydrodynamic pressure reaches 260 MPa and numerical contact occurs on the bearing edges. Mixed lubrication takes place near these areas because of very low values of the film thickness. 4.1 Basic Idea. When performing parametric studies, the input parameters are set to predefined values within a given study range"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002033_icelmach.2014.6960180-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002033_icelmach.2014.6960180-Figure3-1.png",
        "caption": "Fig. 3. No-load flux distribution by FEM magneto-static simulation.",
        "texts": [
            " 2, V0 indicates the no-load rms voltage, I0 the no-load rms current and V0\u2019 the rms voltage drop across the magnetizing reactance and the portion of the stator reactance that excludes end-coil effects. As regards the magnetizing inductance Xm, it is computed as a function of the magnetizing current 0I by running magneto-static FEA simulations where the magnetizing current 2/3/02 I is imposed in one stator phase (a), while in the other two phases (b, c) a current equal to 2/3/02 I is imposed. An example magneto-static FEA simulation output is provided in Fig. 3. From the FEA simulation, the profile is obtained for the air-gap normal flux density and, by Fourier series decomposition, the first spatial harmonic (fundamental) of such wave is deduced. The peak value of the fundamental flux density waveform is called Bn1st. Hence the average airgap flux density is computed as: 2 1 stnn BB med (1) and the flux per pole at the air gap is: cvpnp kLB med (2) where p is the pole span, L the core length, kcv the laminations fill factor. The rms value of the flux linkage for a stator phase due to the air-gap flux fundamental is: 2 1 wspt kN (3) where Ns is the number of series-connected turns per phase and kw is the stator winding factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002646_icaees.2016.7888114-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002646_icaees.2016.7888114-Figure2-1.png",
        "caption": "Figure 2: Experimental apparatus setup for propeller rotation measurement.",
        "texts": [
            " The principle of force measuring is based on the law, the balancing of forces, in which the upward force (lift) created by the motor and propeller combination on one side is balanced by the force on the weighing scale on the other. Figure 3 shows the illustration of the forces acting on the apparatus and motors. In the second part of the experiment, the same motor/propeller combination was installed on a vertical jig to measure its rotation speed. A passive optical tachometer is used to determine the RPM of the motor. Figure 2 shows the experiment setup for propeller speed measurement. B. Electrical Circuit For The Force Measurement Experiment The electrical circuit consists of Lithium polymer (Lipo) battery, servo tester, ESC, power analyzer and DC brushless motor. A LiPo 3S 12V battery rated at 9000 mAh with a 25-50 C discharge capacity is used as the main source of power. A power analyzer is connected in series to this power source and the ESC. The measurements of peak current, voltage and power consumption are taken directly from the power analyzer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000563_aim.2015.7222744-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000563_aim.2015.7222744-Figure2-1.png",
        "caption": "Fig. 2. Schematic representation of the 3-PUU PKM",
        "texts": [
            " Let the x-axis direct along \u2212\u2212\u2192 OB1 and be parallel to x\u2032 axis, and the z and z\u2032 axes be normal to \u2206B1B2B3 and \u2206A1A2A3, respectively. The moving platform possesses only translational motion DOF, therefore, define r = [x, y, z]T as the position vector of the moving platform with respect to the fixed frame O \u2212 xyz. III. INVERSE KINEMATIC ANALYSIS A. Inverse Position Analysis Given the position of the moving platform, inverse position analysis of the 3-PUU PKM refers to determining the linear displacements of sliders. As shown in Fig. 2, the position vector r = [x, y, z] T of point O\u2032 can be expressed by r = bi + qie+ di + lwi \u2212 ai, (i = 1, 2, 3) (1) where ai = [ s cos\u03b2i \u2212s sin\u03b2i 0 ]T (2) bi = [ S cos\u03b2i \u2212S sin\u03b2i 0 ]T (3) di = [ \u2212d cos\u03b2i d sin\u03b2i 0 ]T (4) e = [ 0 0 1 ]T (5) where s, S, qi, \u03b2i, l, e, wi, di, ai and bi denote the radius of the moving platform, the radius of the fixed base, the linear displacement of slider, the angular of point Bi in the coordinate O \u2212 xyz, the length of the strut, the unit vector along the lead screw, the vector along the strut, the vector from a lead screw to the center point of universal joint Ci, the position vector of point Ai in the coordinate O\u2032 \u2212x\u2032y\u2032z\u2032 and the position vector of point Bi in the coordinate O\u2212xyz, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000720_j.cja.2015.08.010-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000720_j.cja.2015.08.010-Figure1-1.png",
        "caption": "Figure 1 Schematic diagram of ECR helicopter.",
        "texts": [
            " 23 24 25 26 27 The concept of swashplateless rotor was brought forward at the beginning of 21th century.1 This kind of rotor system applies blade pitch inputs via purely active flap control or 28 29 30 31 32 33 34 35 36 blade root indexing and active flap for cyclic control instead of traditional swashplate mechanism.2,3 Electrically controlled rotor (ECR) is one kind of swashplateless rotor. For an ECR helicopter, primary flight control is provided by applying blade pitch inputs via integrated active trailing edge flap. Fig. 1 shows the schematic diagram of the ECR helicopter. ECR has many new features,4,5 such as simplified mechanical control system, reduced parasite drag caused by the hub, redundant design for more reliable and safer control system, etc. More importantly, ECR can achieve independent blade pitch control absolutely without the restraint of swashplate and any form of the control input can be achieved theoretically. Then taking this advantage appropriately, it is potential to achieve ECR performance enhancement using active control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002626_iccas.2013.6704133-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002626_iccas.2013.6704133-Figure2-1.png",
        "caption": "Fig. 2 Measurement between a mobile robot and a line landmark.",
        "texts": [
            " In this paper, we generate an observability matrix for line landmarks and point landmarks, and we analyze the degree of observability according to the type or state of the features using the condition number. The state transi tion vector f(x, u) is used as the vector (5). x = f(x, u) \ufffd [vcos(1jJ) vsin(1/)) w]T (5) Using this state transition vector, we derive the analytic solution of the condition number of the observability ma trix based on the line landmarks in Section 3. In the relationship between the robot and the line land mark shown in Fig. 2, the measurement model is ex pressed as (6). z = h(x) = [\ufffd] = [Ir -xco\ufffd(\ufffd ; YSin(b)l ] (6) We can obtain the observability matrix relative to the line measurement model by substituting this model into the observability matrix given in Section 2. Several re searchers have already proved that more than two land marks must be present to estimate the robot's state in 2D space so we investigate the condition number and the phenomena in a situation containing two line landmarks. When two line landmarks exist, the observability matrix is as shown below"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001532_0954406214536700-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001532_0954406214536700-Figure1-1.png",
        "caption": "Figure 1. Section of gear billet on end-face.",
        "texts": [
            " In this article, according to the meshing theory of non-circular gears, two fundamental linkage-models of external non-circular helical gears shaping are built. Eight adoptable shaping strategies and their practical linkage-models are developed according to a comprehensive study. An optimal strategy and its linkage-model have been singled out progressively according to a kinematic analysis and a kinetic analysis, which are illustrated with a virtual shaping and a shaping experiment. Linkage-models for external non-circular helical gears Shaping motions and coordinate systems As shown in Figure 1, a gear billet of non-circular gear is driven by a stage. A workpiece coordinate system Sc\u00f0oc xcyczc\u00de, fixed with machine, is built with its origin locating in the gyration center of gear billet. The xc-axis passes through the gyration center of shaper cutter, and the zc-axis is coincident with the spindle of the stage. A cutting tool coordinate system Sb(ob xbybzb), revolving synchronously with the shaper cutter, is built with its origin locating in the gyration center of the cutter. At the beginning, The xb-axis and the yb-axis are independently parallel to the xc-axis and the yc-axis",
            " The pitch curve of gear billet is going to be always tangent to the pitch circle of shaper cutter, and in continuous pure rolling with it, which constitutes a generating motion. The gear billet moves along the x-axis (vx) in plane, which generates a non-circular pitch curve. As for helical gears, additional rotations !c or !b is essential, which can generate helixes of gears. Moreover, the shaper cutter should move up and down along the z-axis (vz), namely primary motion. Linkage-models in plane. As shown in Figure 1, in order to display shaping process in plane, the shaping motions can be analyzed by taking a section of gear billet on end-face as the stationary frame of reference. After shaping for some time, the location of shaper cutter moves from ob to o0b. According to the meshing theory of non-circular gears,19 the pitch circle of shaper cutter is pure rolling along the pitch curve of gear billet. The arc-length AT on the pitch curve of gear billet should equal to the meshing arc-length A0T on the pitch circle of shaper cutter",
            "20 l \u00bc o0boc \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe r2bj \u00fe 2rrbj sin q \u00f03\u00de \u00bc arcsin rbj cos l \u00f04\u00de From equation (3), dl d \u00bc r dr d \u00fe rbj dr d sin \u00fe rrbj d d cos l \u00f05\u00de Hence, vx \u00bc dl dt \u00bc r dr d \u00fe rbj dr d sin \u00fe rrbj d d cos l !p \u00f06\u00de where !p equals to d /dt. The arc-length AT, being equal to A0T, is as follows.19 s \u00bc Z 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe dr=d \u00f0 \u00de 2 q d \u00f07\u00de \u00bc \u00bc s=rbj \u00f08\u00de !b !p \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe dr=d \u00f0 \u00de 2 q rbj \u00f09\u00de where !b equals to d\u2019/dt. As shown in Figure 1, \u00bc \u00fe p/2. From equation (8), \u00bc s=rbj \u00fe \u00fe =2 \u00f010\u00de Taking the derivative of equation (10) and simplifying it, !r !p \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe dr=d \u00f0 \u00de 2 q rbj \u00fe dr=d \u00f0 \u00de 2 r d2r=d 2 r2 \u00fe dr=d \u00f0 \u00de 2 \u00fe 1 \u00f011\u00de where !r equals to d /dt. From equation (4), !c !p \u00bc lrbj sin d d \u00fe rbj cos dl d l2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 rbj cos l 2q \u00fe 1 \u00f012\u00de where !c equals to d /dt. Case (i). Additional rotation on gear billet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003540_s1064230720050081-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003540_s1064230720050081-Figure1-1.png",
        "caption": "Fig. 1. Three-link hinge mechanism as model of person swinging on swing.",
        "texts": [
            " It is shown that the transition to oscillation with a constant amplitude is implemented. If the friction at the suspension point is relatively small, then the oscillations turn into a rotation and the swing rotates around the suspension point like a gymnast performing the \u201csun wheel\u201d exercise on the bar. Thus, in this paper, the appearance of swing oscillations, together with a swing\u2019s circular rotation is interpreted not as the result of parametric excitement but as the result of a purposeful feedback control. Figure 1 shows a pattern of the considered f lat three-link model. Link HP models a human body, together with the arms and head; link PK models two human hips rigidly connected to the swing; and link KF models both shins. Hinge P models both human hip joints, hinge K models both knee joints, and O is the point of the swing\u2019s suspension. Points , , and are the centers of mass for links HP, PK, and KF, respectively. The swing consists of rod OCh and the seat perpendicular to it and rigidly fastened to rod OCh. Link PK (hips) is located on the seat and is rigidly fastened to it (in Fig. 1, the seat is not shown). Thus, link PK is orthogonal to rod OCh and makes up one solid with it. We do not take into account the mass of the swing (i.e., of rod OCh together with the seat): we consider this mass to be negligibly small compared to the mass of a man (the three-link mechanism, HPKF). The system under consideration has three degrees of freedom and, accordingly, three generalized coordinates: angle \u03d5 between the vertical and the swing (rod OCh), together with angles \u03b1 and \u03b2 of the deviation of links KF and HP from the straight line parallel to rod OCh. Hence, the angle between link KF (shins) and the vertical is (\u03d5 + \u03b1), while the angle between link HP (body) and the vertical is (\u03d5 + \u03b2). The counterclockwise direction is chosen as positive when counting the angles. bC hC sC JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 59 No. 5 2020 In hinge K (knee joints), the control moment L is applied; in hinge P (hip joints), the control moment M is applied (see Fig. 1). Each of these moments is considered to be limited in magnitude: (1.1) Moments L and M can rotate link KF (shins) and link HP (body) relative to link PK (hips), respectively. Assume that viscous friction forces act at the suspension point of swing O and their moment is proportional to the angular speed ; here, \u03bc = const. Due to the anatomical properties of a person, the angles of the knee and hip joints can change only to a limited extent. These restrictions are due, in particular, to the ligaments and tendons, and for the body, also the arms",
            " We will model the appropriate impacts as moments of one-sided spiral springs with high stiffness in the K and P hinges. A model of this kind is considered in [25]. When angles and reach the critical values \u03b1min and \u03b1max, \u03b2min and \u03b2max, in hinges K and P, the moments of the spring forces LS and MS arise; they are intended to return angles \u03b1 and \u03b2 to ranges in which they can change freely: (1.2) The ranges in which linear springs are not affected are denoted as follows: (1.3) It is natural to consider that We derive the motion equations of the system shown in Fig. 1 using Lagrange\u2019s second method. For this we need to write out expressions for the kinetic and potential energy, as well as for the virtual work of moments L and M. The expression for kinetic energy T is presented as follows: (2.1) Here, (2.2) ( )\u2264 \u2264 =0 0 0 0, , const .L L M M L M \u03bc\u03d5 \u03d5 \u03b1 \u03b2 \u03b1 < \u03b1 < \u03b1 \u03b2 \u03b2 < \u03b2   = \u03b1 \u2212 \u03b1 \u03b1 \u2265 \u03b1 = \u03b2 \u2212 \u03b2 \u03b2 \u2265 \u03b2    \u03b2 \u2212 \u03b2 \u03b2 \u2264 \u03b2\u03b1 \u2212 \u03b1 \u03b1 \u2264 \u03b1  min max min max max max max max min minmin min 0 for , 0 for < , ( ) for , ( ) for , ( ) for .( ) for , S S S S SS L k M n nk \u03b1 \u2264 \u03b1 \u2264 \u03b2 \u2264 \u03b2 \u2264\u03b1 \u03b2min mimax maxn\u00a0 , \u00a0 ",
            "11) have the dimension of the moment of inertia (see (2.2)). In addition to the control moments L and M, Eqs. (2.7) and (2.8) also take into account the moments LS and MS (see formulas (1.2)). In order to build the control (in the form of feedback) in the three-link swing model, first, we consider an auxiliary (simplified) swing model that contains only two links. By exploring this simpler model, it is possible to build the quasi-optimal control, which is then used to control the more complex three-link model shown in Fig. 1. Figure 2 shows a relatively simple model of a swing with a person on it. As mentioned above, such a model has been considered in many works. The fixed hinge O is the suspension point of the swing OC, which is considered to be weightless, and l is its length. As in the three-link model, we denote by \u03d5 the angle of the swing\u2019s deviation from the vertical. A person is modeled by the absolutely solid homogeneous rod HF, whose center of mass C is hinged to the free end of the swing. We denote by \u03b8 the angle between the continuation of the swing OC and the rod HF",
            "2) given above, the inequalities and or the values \u03b1min and \u03b1max, together with \u03b2min and \u03b2max. \u2212 < \u2212 = \u2264  > \u03b2 \u2265= \u2212 \u03b2 \u2212 \u03b2 \u2212 \u03b2 \u03b2 = \u03b2 < 0 1 0 1 1 0 0 1 0 max 1 1 prog 2 prog min for , for , for , for \u03c6cos\u03c6 0, ( ) , for \u03c6cos\u03c6 0. M M M M M M M M M M M n n \u03c6cos\u03c6 \u03b1 \u03b2 >2 0k >2 0n \u2265 \u03c6cos\u03c6 0 \u03b1max \u03b2max \u03b1 \u03b2 < \u03c6cos\u03c6 0 \u03b1min \u03b2min \u03d5 \u03d5 \u03c6cos\u03c6 \u03d5 \u03d5 \u2265 cos 0 \u03d5 \u03d5 < cos 0 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 59 No. 5 2020 \u03d5 The numerical studies of the equation of motion (2.6)\u2013(2.8) of the three-link hinge mechanism shown in Fig. 1, with moments (1.2) and control (5.1) and (5.2), are conducted for the following anthropomorphic parameter values of this mechanism, which are taken from [27]: The parameters of control laws (5.1) and (5.2) are The values of the limits on the control moment are chosen using the experimental data presented in [28]. Figure 8 shows for a relatively high value of the viscous friction factor \u03bc = 10 N m s, the swinging process characterized by the zero initial values of the angle of deviation and the angular speed of the swing",
            " In this mode, the full turning time of the swing is T \u2248 2 s. The periodic rotation mode is orbitally asymptotically stable. Thus, if the friction factor is suficiently small, then with control (5.1) and (5.2), the swing (after some oscillations with increasing amplitude) comes to the stationary motion-mode, which represents a rotation around the suspension point. In the rotation presented in Fig. 11, the angular speed and the person \u201cdoes a sun wheel\u201d on the swing; here, the person rotates face frontward, i.e., counterclockwise (see Fig. 1). Depending on the initial conditions, it is also possible to enter the rotation mode clockwise (dashed curve in Fig. 7). The function \u03d5(t) has discontinuities if it is considered in the band of the phase plane (see Fig. 11). However, if the band is \u201crolled up\u201d into a cylinder by combining the horizontal straight lines and \u03d5 = \u03c0, then instead of a phase plane we obtain a phase cylinder, on which the graph of function \u03d5(t) is continuous and spiral. Figure 10 shows not only two solid curves that are related to the process of pumping the swing and entering the rotation mode but also two dashed curves",
            " The numerical study shows that due to the rather high stiffness of the springs described by moments and , there is only a minor violation of conditions (1.3) in the process of the pumping of the swing and in the steady modes. Note that the magnitude of angles and do not exceed the value of \u03c0/2 any time during the motion. In the absence of moments and , this restriction is not always fulfilled. CONCLUSIONS There is no external control moment at the suspension point of a swing; i.e., the corresponding degree of freedom is passive. Thus, in the mechanical model of the swing (Fig. 1), there are three degrees of freedom and only two control moments; i.e., the system is underactuated. The person on the swing in the seated position (as well as in the standing position) can apply only internal forces, namely, moments in the joints, and move some parts (links) of his body relative to others. However, with such movements, the configuration of the person changes, together with the location of his center of mass: his moment of inertia relative to the suspension point of the swing also changes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002528_detc2016-59732-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002528_detc2016-59732-Figure3-1.png",
        "caption": "FIGURE 3: Longitudinal forces on Rav4EV",
        "texts": [
            " While the GPS was used to record the vehicle speed and acceleration, the vehicle CAN was used to record the triggers for acceleration and braking in order to evaluate the vehicle response during acceleration and braking. The data recorded by the VMS, GPS and vehicle CAN was integrated with the help of a CAN integration device from \u2018Vector Informatik GmbH\u2019 so as to have a common time stamp for the collected data. The system architecture for integrating the three devices using the Vector is shown in Fig. 2. The force balance equation for the longitudinal model of the vehicle shown in Fig. 3: Mx\u0308 = Fx f +Fxr \u2212 ( Fr f +Frr ) \u2212Fd (1) where Fx f and Fxr are the longitudinal forces caused due to fric- tion acting on the front and rear wheels, Fr f and Frr are rolling resistance forces on the front and rear wheels and Fd is the aerodynamic drag force on the vehicle assumed to act at the center of gravity (CG). Also Fz f is the sum of normal forces on the front left and right wheels and Fzr is the sum of normal forces on the rear left and right wheels. The location of the CG from the front 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings",
            "org/about-asme/terms-of-use be ignored [6] which reduces the final coast down time T to: T = [ 2M2 \u03c1CdA f Rx ] 1 2 tan\u22121 [ Vi( \u03c1CdA f 2Rx ) 1 2 ] (6) From equation (4) and (6) we get: t T = 1\u2212 tan\u22121 [ Vx( \u03c1CdA f 2Rx ) 1 2 ] tan\u22121 [ Vi( \u03c1CdA f 2Rx ) 1 2 ] (7) Solving for velocity Vx from equation (7) yields: Vx = ( 2R \u03c1CdA f ) 1 2 tan {( 1\u2212 t T ) tan\u22121 [ Vi ( \u03c1CdA f 2Rx ) 1 2 ]} (8) Simplifying further, Vx Vi = 1 \u03b2 tan [( 1\u2212 t T ) tan\u22121 \u03b2 ] (9) where \u03b2 =Vi ( \u03c1CdA f 2Rx ) 1 2 (10) Further, from equation (5) and equation (9), the coefficient of drag Cd is obtained as: Cd = 2M\u03b2 tan\u22121\u03b2 ViT \u03c1A f (11) and coefficient of rolling resistance fr is obtained as: fr = Vitan\u22121\u03b2 \u03b2T g (12) Using equation (9), the parameter \u03b2 is estimated through a least square curve fit by plotting non-dimensional velocity Vx Vi against non-dimensional time t T in Fig. 8. Data was collected over 5 runs of the test track. It can be seen from the plot in Fig 9 that the average coefficient of drag (Cd) is calculated as 0.3083 which FIGURE 8: Plot of Vx V0 against t T to estimate \u03b2 is close to 0.3 specified in the Toyota technical manual. Also, the average coefficient of rolling resistance ( fr) is estimated as 0.01012. The normal forces on the front and rear wheels Fz f and Fzr as shown in Fig. 3 are calculated by taking moments about the contact point of the rear and front tires respectively and expressed as: Fz f = MgLr L \u2212 Mhx\u0308 L \u2212 Fdhd L Fzr = MgL f L + Mhx\u0308 L + Fdhd L (13) The height of CG is estimated using equations (13) above as the forces Fz f , Fzr and M are measured using the WPS sensors (VMS 5 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90687/ on 02/23/2017 Terms of Use: http://www.asme"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000355_b978-0-12-391927-4.10019-2-Figure19.7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000355_b978-0-12-391927-4.10019-2-Figure19.7-1.png",
        "caption": "FIGURE 19.7 Three sources of line tension, l (in units of N): (a) The edge of a liquid drop on a surface where the surface energies gLV and gSL near the three-phase line are changed due to the interaction across the thin liquid film in that region. For q< 90 the line tension is positive (l> 0); for q> 90 it is negative (l< 0). (b) The edge of a monolayer in the solid or crystalline state (see Worked Example 19.2). Monolayers and bilayers in the fluid state can deform to relax the high edge energy \u2013 effectively reducing the length \u2018 in Eq. 19.22 for the same volume v, and these usually have much lower line tensions. (c) The edge of the charged or zwitterionic (dipolar) headgroup region of a surfactant or lipid monolayer where the electric field energy is different from the uniform field within the monolayer, giving rise to a negative line tension whose value depends on the domain radius R (Eq. 19.21).",
        "texts": [
            " The concept of the line tension was introduced in Eq. (19.8) as a parameter that quantifies the (usually unfavorable) energy of the edge of a 2D aggregate or a 3-phase boundary line. Line tensions arise when the edge of a monolayer or droplet lens contributes an additional positive or negative energy to the total surface energy. The line tension l is in units of energy per unit length (J m 1 or N) rather than energy per unit area and can be due to a number of different effects including van der Waals, hydrophobic, and electrostatic forces (Figure 19.7). n n n Worked Example 19.2 Question: Estimate the line tension of a surfactant or lipid monolayer on the water-air interface (Figure 19.7b) assuming that it exposes a hydrocarbon edge of thickness \u2018 \u00bc 2 nm? Answer: The line tension is l \u00bc edge energy per unit length \u00bc \u00f0surface energy of edge; J m 2\u00de \u00f0area of unit length of edge; m2\u00de=\u00f0unit length; m\u00de\u00bc \u2018g \u00bc \u00f027 10 3\u00de \u00f02:0 10 9 1:0\u00de=\u00f01:0\u00de \u00bc 5:4 10 11 J m 1 \u00bc 5:4 10 11N. Typically measured values are 10 11 10 10 N for monolayer domains in the gel, crystalline, or solid state, as schematized in Figure 19.7b, and 10 14 10 12 N for fluid state monolayers and bilayers (Veatch and Keller, 2003). n n n Another contribution to l comes from the charges present in the headgroups of surfactant or lipid molecules that can be modeled as a capacitor (Figure 19.7b and c) consisting of charges q, separated by a \u201cdipole\u201d distance D, occupying an area a0 corresponding to a surface density of G \u00bc 1/a0. The dipole moment u can therefore be expressed as u \u00bc qD and the dipole moment density is uG. The energy per unit area of a capacitor with plates of charge density s, separated by a distance D in a medium of dielectric constant 3 was previously given as s2D=2303 per unit area (see Figure 3.2 and Eq. 3.7). For a circular monolayer of radius R, the total energy is therefore expected to be pR2s2D=2303 \u00bc pR2\u00f0uG\u00de2=2303D: However, this is to ignore the additional term from the edges where the electric field is distorted, as shown in Figure 19.7c. For D \u00ab R the total electrostatic energy is6 pR2\u00f0uG\u00de2 2303D R\u00f0uG\u00de2 2303 log 16pR eD \u00bc pR 2\u00f0uG\u00de2 2303D 1 D pR log 16pR eD : (19.20) 6In these equations e \u00bc 2.718. The correction term to the capacitance C of a condenser is known as Kirchhoff\u2019s formula, where the total energy is Q2/2C where Q is the total charge on each plate. Chapter 19 \u2022 Thermodynamic Principles of Self-Assembly 523 The first term in Eq. (19.20) is proportional to the area and therefore to the number of molecules in the monolayer, N \u00bc pR2G, while the second term is proportional to the perimeter, 2pR, and therefore corresponds to a line tension contribution of lel \u00bc \u00f0uG\u00de2 4p303 log 16pR eD : (19",
            "5 1018 m 2, then uG \u00bc 2.003 10 10 C m 1, and we obtain a domain radius of about 75 nm using Eq. (19.24a) and 2,600 nm \u00bc 2.6 mm using McConnell\u2019s equation. Typically observed sizes of surfactant domains at the water-air interface range from the very small (nano sized) to the very large (many microns), but whether they are true equilibrium structures has yet to be established. These domains tend not to coalesce because of the long-range dipole-dipole repulsion between them (see electric field lines in Figure 19.7c). Their growth occurs via slow Ostwald ripening of surfactant diffusing on the water surface (see Figure 19.6). 2D micelles occur in both monolayers and bilayers, where they are also referred to as domains, patches and rafts. They are discussed further in Section 20.10 and Chapter 21. 7The dielectric constant of the hydrophilic headgroup region is unknown but is expected to be less than the value for bulk water. When the area of a soluble monolayer is compressed, the surfactants go into the solution, thereby keeping the surface coverage and pressure (tension) unchanged"
        ],
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    },
    {
        "image_filename": "designv11_30_0002557_jfm.2016.816-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002557_jfm.2016.816-Figure1-1.png",
        "caption": "FIGURE 1. (Colour online) Schematic of the dimensionless problem, indicating the particle autophoretic motion.",
        "texts": [
            " For convenience we calculate the flow about a stationary particle. Once the resulting hydrodynamic force on such a particle is determined the associated velocity of a comparable freely suspended particle is readily obtained using the well-known mobility expressions for the motion of a sphere perpendicular to a solid wall. We employ a dimensionless notation where all length variables are normalised by a. We employ the cylindrical coordinates (\u03c1, \u03c6, z) with corresponding unit vectors (e\u0302\u03c1, e\u0302\u03c6, e\u0302z), see figure 1. The plane z = 0 coincides with the fluid\u2013wall interface while the z-axis, which points into the fluid, passes through the particle centre. The azimuthal angle \u03c6 is degenerate in the axisymmetric problem considered. http:/www.cambridge.org/core/terms. http://dx.doi.org/10.1017/jfm.2016.816 Downloaded from http:/www.cambridge.org/core. University of Memphis, on 28 Dec 2016 at 10:32:25, subject to the Cambridge Core terms of use, available at We employ the macroscale description of Michelin & Lauga (2014), neglecting however solute advection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001181_j.jeurceramsoc.2014.01.001-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001181_j.jeurceramsoc.2014.01.001-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of gear.",
        "texts": [
            " hile the content should be as little as possible for the purpose f fully discharging binder in the following degreasing process nd eventually increasing the body density. Consequently, 8 wt% as chosen as the optimal weight content of E06 in the composite owder. .2. SLP/CIP/SSS experiment The alumina composite powder was selective laser processed n the HRP-IIA type rapid prototyping system (Huazhong Uniersity of Science and Technology). The optimal SLP processing \u22640.03 \u22653.97 \u226595 arameters obtained from density orthogonal experiments are resented in Table 2, by which gear components with same size ere formed for repeatability. As shown in Fig. 2, the dimenions of the gears are listed as 53.88 mm in D, 41.38 mm in T, 0.8 mm in H. Firstly, the SLPed gear components were coated with natural ubber latex with a thickness of approximately 1 mm. Secondly, he components were CIPed under the hydrostatic pressure f 320 MPa. And then the CIPed components underwent the egreasing process based on the thermal gravity (TG) of the 1856 Z. Wang et al. / Journal of the European Ceramic Society 34 (2014) 1853\u20131863 Table 2 Optimal SLP processing parameters by density orthogonal tests"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003511_j.triboint.2020.106696-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003511_j.triboint.2020.106696-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram of structural parameters of brush seal.",
        "texts": [
            " That is to say FN is a function of time. When the Generally linear beam theory is applied to calculate FN(t) caused by the contacting between bristles and rotor which is usually a small deflection case. The diagram of deflection and force acting on a single bristle is shown in Fig. 2. When operating, the rotor will deform due to centrifugal force and thermal stress, and the bristle will deflect following the deformation. The calculation-related structural parameters of the brush seal are shown in Fig. 3. At the root of the bristle pack, the bristles are welded and can be considered to be stacked closely, and the thickness is Bu, mm. At the tip, the bristles are free and the thickness is Bd, mm. When working, the pressure drop will compress the bristle pack, and the tip thickness will decrease, so, the true tip thickness is between Bu and Bd. The normal contact load for the bristle pack can be considered as the linear superposition of the loads of all the bristles. According to Demiroglu [15], the normal tip force of a single bristle can be written as following, FNs(t)= 3\u03c0 64 Ed4 l3sin 2 \u03b2 \u03b4r(t) (12) where subscript s refers to a single bristle; E is Young\u2019s modulus of bristle material, MPa; d and l are the diameter of a bristle and its free length, respectively, mm; \u03b2 is the lay angle, degree; and \u03b4r(t) is the radial deflection of the bristle tip, mm",
            " In this paper, Di = 110 mm, and the maximum error of the 3D scanner is 0.0107 mm, so the measurement values are accurate to 0.01 mm. Three-Coordinate Measuring Machine (CMM) shown in Fig. 12 is used to measure the Do. The maximum error of CMM is also related to the nominal size. The formula to calculate the maximum error is (0.6+ Do/600) \u03bcm. In this paper, Do = 110 mm, then the maximum error of CMM is 0.783 \u03bcm, so the measurement values are accurate to 0.001 mm. A typical brush seal was used for this study. The brush seal specimen structure is shown in Fig. 3 and its details are listed in Table 2. The bristle material Inconel 178 is a kind of nickel base superalloy, and the hardness is 346\u2013450 HBS, the compressive elastic modulus is 203 GPa, the bending stiffness is 2.39e-7 N\u22c5m2. A rotor was designed made of 38CrMoAl to match the brush seal. The thickness of coating 0.15 mm, and the surface roughness of the coating is 0.02 \u03bcm. The crucial initial geometric tolerances of the brush seal and rotor directly decided the initial interference and the accuracy of the test, which have been measured before testing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001548_iciea.2015.7334299-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001548_iciea.2015.7334299-Figure8-1.png",
        "caption": "Fig. 8. Flux density clouds figure",
        "texts": [
            " Equation(11) shows eddy current loss is proportional to the square of alternating frequency and the square of maximum flux density , is inverse proportional to the resistivity , and related to the structural parameters of PM. B. Simulation model A Two-dimensional model of FSCW-PMSM with 24-slot 16-pole is built by using FEM. On the basis of cycle symmetry, the whole model is divided into 1/8, and it is shown in Fig.6. Considering different area of the motor and the influence of skin effect, the grid subdivision shown in Fig.7. Fig.8 and Fig.9 represent the distribution of the flux density and the magnetic field lines on the motor model. The PM eddy current loss under rated-load condition is shown in Fig.10. It fluctuates periodically, because it changes with the change of the stator coil position. 1248 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) C. Eddy current loss of main harmonics The fundamental harmonic of air gap field has the same speed as rotor, so it does not produce eddy current on PMs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003627_rpj-01-2020-0009-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003627_rpj-01-2020-0009-Figure1-1.png",
        "caption": "Figure 1 The designed coupon for surface finish investigation",
        "texts": [
            " AM via laser melting (specifically, SLM) was selected because the process has an inherently low surface roughness and also allows the printing of 17\u20134 PH stainless steel, which is a standard alloy used in surgical instruments. The mass surface finishing processes developed herein aim to reduce the surface roughness of AM parts to the level of an equivalent machined part while maintaining part integrity. The coupon incorporated a variety of features commonly found on surgical instruments was developed (Figure 1a). Features included multiple surface orientations, internal and external radii and a through-hole. Support structures were left intact on Surface finishing Alasdair Soja, Jun Li, Seamus Tredinnick and TimWoodfield Rapid Prototyping Journal one because these are generally unavoidable for complex printed parts. The final dimensions of the part were chosen to ensure the representative surfaces were large enough to accommodate stylus profilometrymeasurements. Surfaces A, B and C were horizontal (inclination angle: 0\u00b0), inclined (30\u00b0) and vertical (90\u00b0) to the build platform",
            " The samples were manufactured by Linear AMS (Livona, MI, USA) using the SLM process on an EOS-INT M 280 (EOS GmbH, Germany), with a 400W laser and EOS 17\u20134 PH stainless steel powder. The composition of the powder was as follows: Cr: 15\u201317.50%, Ni: 3\u20135%, Cu: 3\u20135%, Si: 0\u20131%,Mn: 0\u2013 1%, P: 0\u20130.04%, S: 0\u20130.03%, C: 0\u20130.07%, Nb 1 Ta 0.15\u2013 0.45%, with the particles having a range in size of 16\u201363mm (Stirling, 2015). Parts were wire-cut from the build platform by electrical discharge machining (EDM). The support structures were snapped off and the remnants reduced using a file. Approximately 0.3mm of the material was removed from Surface E (Figure 1a) using an 8mm end mill. The final asbuilt samples ready for testing are shown in Figure 1(b). Parts were processed as-built or heat treated according to the specifications of the manufacturer of the alloy powder. Briefly, heat-treated parts were solution anneal per UNS17400, and then age hardened to H900 condition. This heat treatment schedule resulted in amaterial hardness of 42.86 0.3Hardness Rockwell C (HRC), increased from the as-built hardness of 236 1.6HRC. Surface roughness was characterised using a Veeco Dektak 150 stylus profilometer with a 2 mm radius stylus. Twodimensional (2D) surface profiles were levelled then split into roughness and waviness using a phase-correct Gaussian filter with a cut-off wavelength of 800 mm (determined by trial measurements), according to ISO11562 (International Organisation for Standardisation, 1996a)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002223_978-3-319-22894-5_19-Figure19.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002223_978-3-319-22894-5_19-Figure19.2-1.png",
        "caption": "Fig. 19.2 Integration of polysiloxane-based side-chain nematic LCE in Microsystems. (a) Preparation of the silicon frame together with the copper electrical circuit and integration of the LCEs with the gold heating wires and the silicone joints before removal of the silicon supporting structures; two pictures showing the mechanical actuation of the LCE microgripper before and after contraction of both LCE films when heating disorder is produced; contraction of the LCE films as function of time at different voltage rates; and (b) integration of the LCE into the microchamber before sealing and the obtaining of the microvalve; sequence of pictures showing the positioning of the LCE film in (a) the nematic phase when the microvalve is fully open, (b) when approaching the isotropic temperature and the microvalve starts closing, and (c) when fully isotropic and the microvalve is closed (Note: the blue arrows show the liquid flow direction); electrical power, temperature of the microchamber and differential pressure between the two openings as function of time when closing the microvalve\u2014the letters correspond to the pictures described before",
        "texts": [
            " This concept was firstly reported by Bru\u0308ndel et al. (2004) who demonstrated for the first time the possibility to integrate LCE materials into MEMS/MOEMS systems, and the compatibility of such materials with some of the manufacturing processing techniques in microelectronics. Following this principle, Sa\u0301nchez-Ferrer et al. (2009) introduced the first LCE-driven silicon microsystem: a microgripper based on the thermal actuation of a monodomain LCE film, showing temperaturecontrolled motions. Figure 19.2a shows a schematic representation of the device with the main parts, as well as pictures of the final device and its actuation principle. For the obtaining of the nematic polysiloxane-based LCE material, a side-chain nematic liquid-crystalline polymer (LCP) containing photo-reactive moieties, which allow the crosslinking of the polymer chains, was synthesized and oriented using a magnetic field of 11 T or an electrostatic field of 300 kV \u00b7m 1. The photocrosslinking of the LCP chains using UV light resulted in a monodomain of LCE sample of about 16 mm 4 mm 0",
            " The actuation principle was basically the same than the previous example, but advantage was taken from the expansion of the material in the other two directions perpendicular to the alignment and bending of the LCE in the direction of the liquid flow which allowed the closing of the microvalve upon the nematic-to-isotropic phase transitions of the material. Thus, opening and sealing of the microfluidic channel was achieved when going back and forward from the isotropic to the nematic state. A schematic view of the structure designed, as well as a sequence of images of the microvalve performance and the corresponding actuation plots are depicted in Fig. 19.2b. The volumetric flow of the medium is guided underneath the actuator (Level 3). A small supporting structure on the chip, which is on the same level as the bearing surfaces for the two ends of the actuator (Level 2) prevents buckling in the normal direction. Thus, the deformation of the valve in the main direction cannot be avoided and is compensated by an elevated channel ground (Level 1). Two identical micromachined chips were assembled together face to face to form the microfluidic system including a 0",
            "30 mm LCE actuator in between as a moving valve. One part of the assembled chip contains a copper circuit on its back side for heating, while the backside of the microchip has the electric contacts and a thermoresistor to measure the temperature as function of the applied electric power. For the fabrication of the microvalve, a nematic monodomain of a side-chain LCE was synthesized following the two-step crosslinking process outlined by Finkelmann et al. (2001b). As shown in the sequence of images in Fig. 19.2b, the LCE microvalve sealed the structure upon heating and filling the room in the directions perpendicular to the director and to the liquid flow up to the wall. When the LCE film reached the wall and tension grew, an abrupt buckling of the actuator occurred in the middle of the LCE film closing the microchannel in the direction of the flow. This middle part of the actuator moved to the microchamber blocking the fluid flow and creating an extra pressure due to the self-clamping at the two ends. The shrinkage of the actuator in the parallel direction to the director aided its movement in the microchamber as a result from a reduction of the friction forces between the actuator and the microstructure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-26-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-26-1.png",
        "caption": "Figure 6-26. The reason why a rotation matrix is needed to relate the world frame with the vehicle frame",
        "texts": [
            " Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 276 F F F F m m y z xx y z \u00e9 \u00eb \u00ea \u00ea \u00ea \u00f9 \u00fb \u00fa \u00fa \u00fa = = = \u00e9 \u00eb \u00ea \u00ea \u00ea \u00f9 \u00fb \u00fa \u00fa \u00fa     x When you add the concept of a body in free fall (that is, adding the gravitational component), note that this component only affects the Z axis of the fixed frame. Remember that you are situated in the fixed frame. This way, you must find a relation between the forces of the drone or base frame and those of the fixed frame (the forces are generated by the drone or base frame, but the model is made with the fixed frame or world). See Figure\u00a06-26. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 277 where F F F u xB yB zB = = = 0 0 This happens because the drone must tilt in order to move. You must apply a matrix that projects the unique thrust force of the drone (in its Z body axis) in each of the inertial axes or a fixed frame, through a rotation matrix based on Euler angles (roll, pitch, and yaw. See Bedford\u2019s dynamics). This is the adaptation of Newton\u2019s second law to a rotating frame (Euler\u2019s second law of motion). For the convenience of the reader and for achieving compatibility with the sensors, this is performed in the body frame (the rotation measures taken by the sensors are made with respect to the body or drone frame)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.70-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.70-1.png",
        "caption": "Figure 3.70 Slide loading diagram",
        "texts": [
            " For a one-of-a-kind design, much time can be saved by choosing an appropriate complete assembly Typical specs: Stroke length: up to 2 m Speed: 1.3 m s\u22121 max Repeatability: \u00b10.01 mm, 0.02 mm, 5 \u03bcm Thrust force: 163 kg max. Accuracy: 15 \u03bcm, \u00b10.1 mm/300 mm Max Acc.: 124 m s\u22122 Lead: 2.5, 4, 5, 10, 16, 20, 30 mm/rev Load: 150 kg max. Inertia: Specified with respect to the drive motor shaft 126 Electromechanical Motion Systems: Design and Simulation When selecting a linear slide, an evaluation of the load weight and location with respect to the center of the carriage is important. See Figure 3.70. Both load forces and torques must not exceed suppliers\u2019 design specifications. Note that offset loads and resulting torques must be limited to specified values to assure bearing life. These specifications are with respect to dynamic (acc/dec) values rather than static loads and must therefore be carefully calculated. Also in X\u2013Y and X\u2013Y\u2013Z assemblies, offset loads, for example a horizontal Y-axis mounted on an X-axis will subject the X-axis to a large torque load when the Y-axis (bearing the load) is fully extended"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002050_s10999-016-9338-1-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002050_s10999-016-9338-1-Figure10-1.png",
        "caption": "Fig. 10 Normal and side contact stress isogram of the special plane when the displacement load is 6dc: a asperity normal contact and b asperity side contact",
        "texts": [
            " Compared with the results of two methods, though the normal contact stresses decrease with the increase of the contact radii of asperities, the normal contact stresses along the same contact radius are similar. In addition, the results from the present side contact model of asperities on the gear tooth profiles are more similar to FE model. Suppose that specific plane is across the initial point of plastic deformation and parallels to the bottom of the asperity, the contact stress isogram in the plane is plotted according to FE model. As shown in Fig. 10a, b, the contact stress isogram in the plane is plotted at the asperity normal or side contact state when the displacement load is 6dc. It is shown in the isograms that the contact stress gradually decreases from the center to around. In Fig. 10a, the contact stress isogram in the plane at asperity normal contact state is a circle, which is similar to the elastic\u2013plastic region distribution of the asperity without friction by the present theoretical model in Fig. 2b. Likewise, the isogram at asperity side contact state is a ellipse in Fig. 10b, which is similar to the elastic\u2013plastic region distribution of asperity with friction by the theoretical model in Fig. 3b. The comparative results show that the normal contact stresses of the asperities on the tooth profiles are similar according to the present elastic\u2013plastic asperity contact model with or without the consideration of friction, or to the finite element models for asperity normal and side contact analysis. The results also indicate that the present theoretical model is well applied to elastic\u2013plastic asperity contact analysis on the rough tooth profiles of a spur gear drive"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002194_1464420716660126-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002194_1464420716660126-Figure1-1.png",
        "caption": "Figure 1. Schematic illustration of experimental setup.",
        "texts": [
            " The data are shown in Table 3, which is designed based on the L9 (33) Taguchi orthogonal array.23 The substrate 1045 with length 100mm and diameter 20mm is clamped on the workstation. Experiment is implemented by a CO2 laser (model Han\u2019s Laser 6000) with spot size of 3mm and automatic feeding. The layer is sintered by radiation heat of the laser as multiple-layer coatings with thickness range of 2.5\u20133.2mm. Step 4: After laser cladding, lateral surfaces of the cladding layer of each substrate are cross sectioned to height 2mm for metallographic analyses as shown in Figure 1. Then the round bar is cut into small slices with total length 20mm by wire electrical discharging machine as shown in Figure 1. Totally, nine specimens are prepared following the same steps for shear strength testing. The shear strength properties of the interfacial bonding are measured. A sleeve with the hole size of 20mm was joined to the round specimen with a diameter of 20mm, height of 2mm, and thickness of 2mm. Shear strength testing is performed at room temperature using a hydraulic universal testing machine (Jinan Precision Testing Equipment Co., Ltd, Model WE-600, max. load 600 kN). A schematic illustration of the experimental facility is shown in Figure 1 and the actual experiment is shown in Figure 2. The facture behavior of the nine specimens after shear strength test is illustrated in Figure 3. The fracture was along the length of the interface direction in two modes: complete or partial fracture. Figure 3 shows that the nickel-based cladding layer in specimen 1\u20133 fell off completely, while the iron-based cladding layer in specimens 4\u20139 was partially fractured. Because the bonding strength between the sleeve and nickel-based cladding layer is much stronger Table 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001205_j.compstruc.2013.03.011-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001205_j.compstruc.2013.03.011-Figure2-1.png",
        "caption": "Fig. 2. (a) Elastic linkage elastically clamped at both ends and (b) its primary structure under the compound effects of the external loadings Hi,Vi, Mi and the redundants Rx, Ry, MR.",
        "texts": [
            " Obtaining the solutions for longer linkages is a challenge for numerical mechanics. Therefore, in Section 5 we construct a numerical techniqu e capable to find all the critical equilibrium configurations of the non-linea rly elastic structure for large displacements without the need of knowing the equilibrium paths. Finally, our conclusions are drawn in Section 6. The model we study is compose d of N rigid links of length \u2018 (i = 1, . . . ,N) attached together by frictionless hinges and connected pairwise by non-linea r rotational springs, as shown in Fig. 2 (a). These links are assumed to be infinitely slender and different links are allowed to cross each other. The position of a link is described by the location of its starting point and its angle: ai is the angle of the (i + 1)th link with the horizontal, yi is the vertical distance of the ith hinge from the horizontal axis x, and xi is the horizontal distance of the ith hinge from the vertical axis y. Here, axes x and y form a Descartes-coordinate system with the origin fixed to hinge \u2018\u20180\u2019\u2019. The structure is supported in a statically indetermi nate way such that both ends of the linkage (hinges \u2018\u20180\u2019\u2019 and \u2018\u2018N\u2019\u2019 in Fig. 2 (a)) are attached to fixed walls, and equipped with non-linea r rotational springs, too. The rotations /left and /right of the walls and their distance xN = d and yN = 0 are fixed. The load exerted on a rigid link is modelled by a force acting at the mid point and a couple. The force on the ith link is represented by its horizontal and vertical components, Hi and Vi, respectively , and the couple on the ith link is denoted by Mi (see Fig. 2 (a)). The load can be dependent on the geometry of the structure, therefore, the magnitudes of Hi,Vi, Mi are taken to be functions of ai 1, xi 1, and yi 1, that is, the force and couple on the ith link may depend on the position of that link: Hi = Hi(ai 1,xi 1,yi 1), Vi = Vi(ai 1,xi 1,yi 1), Mi = Mi(ai 1,xi 1,yi 1). Thus the loading is general in the sense that it can be both potential and non-potential. Before continuing the study of equilibrium states of the linkage, we shortly review some important former results",
            " Thus these bifurcation diagrams are not distorted versions of each other. Since we do not restrict our investigation to potential loadings, we use geometrical, equilibrium , and constitutive equations alongside with boundary conditions. The computation is made on a primary structure. It is created such that the linkage is released from the wall at hinge \u2018\u20180\u2019\u2019, thus it becomes statically determinate . The released constrains are substituted by an unknown redundant force with components Rx, Ry, and an unknown redundant moment MR, as shown in Fig. 2 (b). When the primary structure is unloaded, the springs are stress-free, but the shape of the linkage may not be perfectly straight: the initial, stress-free configuration is given by the angles ai (i = 0, . . . ,N 1) of the links from the horizontal . The initial curvature of the linkage at the ith hinge is j0 i \u00bc ai ai 1. Under the compound effects of the external loadings and the redundants , the moment mi arising in the spring at the ith hinge is assumed to be an invertible, non-linear function of the relative rotation: mi \u00bc mi ai ai 1 j0 i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002428_cgncc.2014.7007267-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002428_cgncc.2014.7007267-Figure2-1.png",
        "caption": "Figure 2. Leader-follower formation method",
        "texts": [
            " Robust on-line FDI-ARC system is designed to detect and identify different failures, battle damage and disturbances, and reconfigures the controller accordingly. As shown before, there are many other control systems which can achieve better performances, a simplified control system, which will be illustrated immediately, is constructed to control formation\u2019s geometry and will be used in following research due to its concision and that disturbance can be easily add to this system. In Leader-Follower formation method, the leading UAV (L in Fig. 2) will fly on the basis of settings (path, velocity), and the follower (F in Fig. 2) should just adjust its velocity, altitude and heading according to the leader, so that the follower can keep the relative position to leader, and can change its position as needed[8]. For this purpose, each UAV should have an onboard autopilot. The structure of this control system is illustrated in Fig. 3. The error of follower\u2019s position can be defined in follower\u2019s velocity frame, as shown in Fig. 4. So the problem can be formulated as how to eliminate this error by adjusting the velocity and heading of the follower"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001545_0954406213489925-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001545_0954406213489925-Figure5-1.png",
        "caption": "Figure 5. Five circles on OCT surface through a point Q of plane bitangency.",
        "texts": [
            " Among the four circular generatrices of a given OCT, the two circles with larger radiuses lie on bitangent planes, and therefore are generalized Villarceau circles. If Sinj j> b/a, then the OCT belongs to a special category, which is beyond the scope of this article. These special OCTs look like barrels and their meridians are banana-shaped.13 These OCTs also have four circular generatrices but their bitangent planes do not pass through the OCT center; they have the double at University of Birmingham on April 21, 2015pic.sagepub.comDownloaded from nature of a circular toroid with a zero offset (OCT) and a circular toroid with the non-zero offset. In Figure 5, five circles pass through the point Q in which two circular generatrices of radius b lie on the same bitangent plane. The case a\u00bc b corresponds to a horn OCT; one can have jaj \u00bc j j. The four circular generatrices are congruent and lie on bitangent planes to the surface, as shown in Figure 6. The points of tangency are double points located at the OCT center of symmetry. At the point O that is the horn OCT center of symmetry, there is a revolute cone that is tangent to the surface; it is a conical point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003993_12.2184834-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003993_12.2184834-Figure4-1.png",
        "caption": "Figure 4. Global and local coordinate systems of a mobile robot.",
        "texts": [
            " Targets #6 and #10 define X axis of the robot coordinate system, targets #3 and #13 define Y axis (figure 3). Proc. of SPIE Vol. 9528 95280L-4 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 09/13/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx A state-space model of the Hercules robot was identified using a motion capture system. The state-space dynamic model of the robot is given by: \ud835\udc7f \ud835\udc61 = \ud835\udc68\ud835\udc7f \ud835\udc61 + \ud835\udc69\ud835\udc7c \ud835\udc61 (1) \ud835\udc80 \ud835\udc61 =  \u00a0\ud835\udc6a\ud835\udc7f \ud835\udc61 + \ud835\udc6b\ud835\udc7c(\ud835\udc61) Where output vector Y(t) of the robot consists of longitudinal speed vx and rotational speed \ud835\udf03 (figure 4): Proc. of SPIE Vol. 9528 95280L-5 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 09/13/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx The input vector U(t) consists of average and difference PWM values for left and right wheel motors: \ud835\udc7c \ud835\udc61 =  \u00a0 \ud835\udc4e\ud835\udc51 (3) \ud835\udc4e = \ud835\udc5d! + \ud835\udc5d! 2 \ud835\udc51 = \ud835\udc5d! \u2212 \ud835\udc5d! 2 Where pR, pL \u2013 are PWM values for right and left wheel motors. The model of the robot was identified using two kinds of input signal: chirp signal and white noise signal. Identification using chirp signal Chirp signal was used for estimation of longitudinal dynamics of the robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003978_978-3-319-14705-5_12-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003978_978-3-319-14705-5_12-Figure12-1.png",
        "caption": "Fig. 12 Two views of the mass during circular motion",
        "texts": [
            " 11), lx , ly are the norms of \u21c0 lx , \u21c0 l y respectively, \u21c0 lxi , \u21c0 l yi are the projections of length L on axes xi and yi respectively (shown in Fig. 11), lxi , lyi are the norms of \u21c0 lxi , \u21c0 l yi respectively. From Eqs. (34) and (35) we can obtain { lxi = p mgr (x\u0308 cos \u03b1 \u2212 y\u0308 sin \u03b1) lyi = \u2212 p mgr (x\u0308 sin \u03b1 + y\u0308 cos \u03b1) + f \u00b7r mg (36) Because \u03b1 = \u03c9t , the driving moment of motor 1 can be got as M1 = mg \u21c0 l j + m\u03b2\u0308y L2 = \u2212 p r (x\u0308 sin \u03b1 + y\u0308 cos \u03b1) + m\u03b2\u0308y L2 (37) where, \u21c0 l j is the projection vector of L on axis j, \u03b2y is the angle that the mass deviates from axis k measured on plane o\u2032k j (as shown in Fig. 12). The driving moment of motor 2 is M2 = mg \u21c0 li + m\u03b2\u0308x L2 = p r (x\u0308 cos \u03b1 \u2212 y\u0308 sin \u03b1) + m\u03b2\u0308x L2 (38) where, \u21c0 li is the projection vector of L on axis i, \u03b2x is the angle that the mass deviates from axis k measured on planes o\u2032ki (as shown in Fig. 12). If spherical robot BHQ-1 moves along a circular trajectory with a fixed velocity its acceleration A = v2 R should point to the center of the circular trajectory and the tangential velocity of circular trajectory v = \u03c9R should be a constant. Where, R is the radius of the circular trajectory, is the angle velocity of BHQ-1. The projections of the acceleration A on axes x and y are { Ax = \u2212A cos(\u03c0 \u2212 \u03b1) = \u2212A cos(\u03c0 \u2212 \u03c9t) Ay = \u2212A sin(\u03c0 \u2212 \u03b1) = \u2212A sin(\u03c0 \u2212 \u03c9t) (39) From, { x\u0308 = Ax y\u0308 = Ay , Eqs. (16), (32), (39) we can get { m0 1(t) = p A r sin(\u03c9t) m0 2(t) = p A r cos(\u03c9t) (40) From Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure9-1.png",
        "caption": "Fig. 9 Mesh of the skirt feature (units in lm). The finer mesh at the iris and hair bulge is also visible.",
        "texts": [
            " This contact region sizing is important to capture sliding of the surfaces as the contact was modeled frictionless. Next, remaining sections were meshed with solid tetrahedron element by using the body sizing control. The element size was chosen in such a manner so that a smooth transition occurs between neighboring parts. The only part for which no specific mesh control was assigned was the hair shaft past the hair bulge. This portion of the hair was also modeled using solid tetrahedron elements. The automatic default ANSYS mesh was used for this part. Completed discretized half model is shown in Fig. 9. 4.2.3 Loading and Boundary Conditions. A fixed support was added to the circumferential face of the skirt (Fig. 10). This skirt feature is cuticular tissue that extends out and connects, eventually to the surrounding tissue of the adjacent hairs. The length of this tissue in the insect is quite long as the distance between two hairs is anywhere from 100\u2013300 lm. However, the radius of the tissue in the model is about 20 lm, which is at least five times smaller than the actual distance between hairs",
            "5 lN was applied at a location as shown in Fig. 23. Displacements of the loading point as a function of total number of elements for mesh 1, 2, 3, and 4 are plotted in Fig. 25. It is observed that the element sizing option 4 is adequate; however, further finer mesh sizing was employed in the following regions for final analysis: contact region between socket base and hair base with 0.5 lm; socket base and hair base with 0.75 lm; and hair shaft to just past the bulge 0.75 lm. The final discretized mesh is previously shown in Fig. 9. [1] Thurm, U., Erler, G., Godde, J., Kastrup, H., and Keil, T., 1983, \u201cCilia Special- ized for Mechanoreception,\u201d J. Submicrosc. Cytol., 15(1), pp. 151\u2013155. [2] Keil, T. A., and Steinbrecht, R. A., 1984, \u201cMechanosensitive and Olfactory Sensilla of Insects,\u201d Insect Ultrastructure, Vol. 2, R. C. King and H. Akai, eds., Plenum Press, New York, pp. 477\u2013516. [3] Keil, T. A., 1997, \u201cFunctional Morphology of Insect Mechanoreceptors,\u201d Microsc. Res. Tech., 39(6), pp. 506\u2013531. [4] Palka, J., Levine, R., and Schubiger, M"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002025_j.apor.2016.04.001-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002025_j.apor.2016.04.001-Figure1-1.png",
        "caption": "Fig. 1. Coordinate systems.",
        "texts": [
            " Kinematics J1( , , ) = \u23a1 \u23a3 cos cos \u2212sin sin cos cos \u2212sin m [ u\u0307\u2212 vr + wq \u2212 xg(q2 + r2) + zg(pr + q\u0307) ] = Xu\u0307u\u0307+ Xuu + { (np/6 m [ v\u0307 \u2212 wp + ur + xg(qp + r\u0307) + zg(qr \u2212 p\u0307) ] = Yv\u0307v\u0307 + Yp\u0307p\u0307 + Yr\u0307 r\u0307 + Yvv m [ w\u0307 \u2212 uq + vp + xg(rp \u2212 q\u0307) \u2212 zg(p2 + q2) ] = Zw\u0307w\u0307 + Zq\u0307q\u0307+ Zww + Ixxp\u0307+ ( Izz \u2212 Iyy ) qr + m [ \u2212zg( v\u0307 \u2212 wp + ur) ] = Kv\u0307v\u0307 + Kp\u0307p\u0307+ Kr\u0307 r\u0307 + K +{ (np/60)2d5 p}K0 + K\u0131el (\u0131er \u2212 \u0131el) + KW Iyyq\u0307+ (Ixx \u2212 Izz) rp + m [ zg(u\u0307\u2212 vr + wq) \u2212 xg (w\u0307 \u2212 uq + vp) ] = Mw\u0307 \u2212xgWcos cos \u2212 zgWsin + M\u0131e (\u0131er + \u0131el) + MW Izzr\u0307 + ( Iyy \u2212 Ixx ) pq + m [ xg (v\u0307 \u2212 wp + ur) ] = Nv\u0307v\u0307 + Np\u0307p\u0307+ Nr\u0307 r\u0307 + Nv Fig. 1 shows the coordinate system used in this study, which onsists of the body- and space-fixed coordinates. The origin of the space-fixed coordinate O \u2212 xyz is located at n arbitrary position on a free surface, and the positive direction esearch 58 (2016) 83\u201394 of depth z is downward. The position and orientation of the submerged body can be defined in the space-fixed coordinate. In the figure, u, v, and w are respectively the surge, sway, and heave velocities; p, q, and r are respectively the roll, pitch, and yaw rates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000789_iros.2013.6696909-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000789_iros.2013.6696909-Figure4-1.png",
        "caption": "Fig. 4. Schematics of the propulsion system onboard CephaloBot.",
        "texts": [
            "00 \u00a92013 IEEE 3865 It is worth mentioning that the feedforward control design described in this paper generally applies to all vehicles, yet it especially suits the need for improving maneuvering accuracy on this particular type of underwater vehicle. In the following sections, the feedforward dynamic model is given and the force estimation algorithm is expressed in three-dimensional space. Nonlinear controllers with the feedforward are tested in simulation for control accuracy in potential small scale maneuvers as well as long distance cruising. Experimental tests are performed to vindicate the force estimation algorithm. A coordinate system in the body-fixed reference frame is defined as shown in Fig. 4. For motion in the horizontal plane, there are three degrees of freedom, namely, translational motions along x- and y-directions (surge and sway), and rotational motion about z-axis (yaw). At a time instant t, the vehicle\u2019s velocity is designated as vector \u03bd(t) \u2208 R 3. The earth-fixed reference frame is considered to be inertial, in which the earth-fixed coordinate system is defined. Position and orientation of the vehicle can be described in the earthfixed frame as vector \u03b7(t) \u2208 R 3. The velocity of the vehicle in the earth-fixed reference frame can be obtained by the following transformation: \u03b7\u0307 = J(\u03b7)\u03bd , (1) where J(\u03b7) \u2208 R 3\u00d73 represents the transformation matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure16.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure16.2-1.png",
        "caption": "Fig. 16.2 Schematic of the push rod valve train and the major components of the friction losses (Teodorescu 2010)",
        "texts": [
            " This include equations obtained through regression of numerical studies that correlates the applied load, entrainment velocity, lubricant viscosity and mechanical properties of contact surfaces for different contact geometries and oil film thickness. In spite of various limitations due to simplifying assumptions, such formulae are relatively easy to use that gives relatively accurate and quick results (Teodorescu 2010). The total valve train energy loss is the sum of the friction energy consumed by all its individual components. A typical valve train system consists of cam, tappet, pushrod, rocker-arm, valve and spring. The major losses occurring in the valve train shown in Fig. 16.2 includes losses between cam and tappet (1), between the tappet and its bore (3), in the rocker arm bearing (5), between the valve stem and the valve guide (7) and in the camshaft bearings. The friction forces at the two ends of the push rod (2, 4) and the friction between the valve stem top and the rocker arm end (6) are much lower than in the other components. They appear as a residual term in the overall balance of frictional losses (Teodorescu 2010). The improvement in automobile industry is mainly focused to satisfy customer\u2019s needs like better fuel economy and strict regulations of government for environment security"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure5-1.png",
        "caption": "Fig. 5. Bertrand-like curves.",
        "texts": [
            " Flank data is more useful when presented in terms of transverse profiles. As such, flank data is resampled using involutoid data or rack data due to the \u201cnon-transverse\u201d direction of the constant pressure angle profile. The involutoid concept presented is restricted to constant speed ratio gears and must be generalized in terms of a spatial evolutoid to support involutoids for non-circular hypoid gears. Namely, the unwrapping of a line in the rectifying plane of the evolutoid curve where the amount of unwrap and the distance along the line varies. Illustrated in Fig. 5 are a family of Bertrand-like curves for each tooth flank where the different Bertrand-like curves share a common center of rotation axis. The distance between these Bertrand-like curves depends on the normal pitch p n . The normal pitch was introduced as follows [1] : p n = 2 \u03c0 \u221a u 2 p + w 2 sin 2 \u03b1p N cos ( \u03b3p + \u03c8 p ) (11) where u p Radius of pitch surface at the throat w Axial distance along pitch surface generator \u03b1p Angle of generator with its rotation axis \u03b3p Angle between transverse curve and perpendicular to generator \u03c8 p Spiral angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003708_5.0015728-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003708_5.0015728-Figure7-1.png",
        "caption": "FIGURE 7. Total Deformation Simulation Results",
        "texts": [],
        "surrounding_texts": [
            "The total deformation simulation results show that the maximum value at the distance between crack 1.5 mm is 2.7558 mm at 4.5 mm, while the minimum value is at 3.0 mm at 0.28674 mm. The maximum value at the distance between 2.5 mm crack explained there is a figure 9 with a maximum value of 2.7588 mm and a minimum value at the point of 5 mm that is 0.3065 mm. When an object that is given a continuous load and is still within the limits of elasticity, it will return to its original form [10]. The maximum value obtained is at the top of the specimen, this happens because the beginning of the load given from the top vertical direction is channeled down, then it will return to the top when the suction and disposal process [11]. The bottom of the specimen only gets the thrust obtained from the rotation of the crankshaft, so the minimum value is at the bottom of the specimen. 040011-4 Stress Intensity Factor Simulation Result The results of the simulation of stress intensity factor with a distance between 1.5 mm crack has a maximum value of 684.07 MPa.mm0,5 at 4.5 mm, while the minimum value is at 3.0 mm with a value of 27,349 MPa.mm0,5. The distance between 2.5 mm cracks in figure 9 has a maximum value of 2269 MPa.mm0,5 at point 2.5 mm and has a minimum at point 5.0 mm which is 126.9 MPa.mm0,5. Based on the results that have been obtained, it appears that the direction of the cracks spread from the outside to the inside. This is due to the pressure force obtained during the combustion process in the engine. If you continue to get the force, in areas that experience cracks will continue to increase the stress intensity factor and will continue to propagate as long as it is given a load that can result in fracture 040011-5 on these components. The value of the stress intensity factor is influenced by the geometry, sample, size and location of the crack and the amount of load distributed to the material. Thus, the presence of cracks in the corner which is a critical region is very influential on the value of the stress intensity factor [12]. Based on research [13] that the tendency to get closer to a crack (through crack), the greater the tendency for smelling errors. Significant difference in value indicates that at the crack site the tendency to approach translucent cracks is also greater."
        ]
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure5-1.png",
        "caption": "Fig. 5 Cutter position in the coordinate system of the generating gear",
        "texts": [
            " Its position vector can be represented in the cutter coordinate system St by the following equation: rt\u00f0u; b\u00de \u00bc yl sin b yl cos b zl 1\u00bd T (3) here y l\u00f0 \u00de l u\u00f0 \u00de \u00bc r0 u cos ab z l\u00f0 \u00de l u\u00f0 \u00de \u00bc u sin ab ( y f\u00f0 \u00de l u\u00f0 \u00de \u00bc y0 \u00fe qb cos u z f\u00f0 \u00de l u\u00f0 \u00de \u00bc z0 \u00fe qb sin u ( and y0 \u00bc r0 qb sec\u00f0ab ac\u00de\u00f0cos ab sin ac\u00de z0 \u00bc qb sec\u00f0ab ac\u00de\u00f0cos ab \u00fe sin ac\u00de where u and b are the parameter of the cutter profile and rotation angle of the cutter. Using the same standard planar-generating gear to cut pinion and gear produces a conjugated gear pair. As shown in Fig. 5, the cutter is moved into the position whose reference point M coincides with the mean point of the generating gear, and the surface normals of both cutter and gear are aligned. Because the cutter surface is conical, it produces a lengthwise-crowning tooth (see Fig. 5). The relative position between the cutter, the generating gear, and the work gear is simulated using a virtual machine. The coordinate systems between the cutter and the generating gear on the virtual machine are as described in Fig. 6, which shows three translatory \u00f0Cx;Cy;Cz\u00de and two rotational axes \u00f0ua;ub\u00de for the cutter position. Here, ua is the cutter tilt angle and ub is the spiral angle. The coordinate systems St\u00f0xt; yt; zt\u00de and Sc\u00f0xc; yc; zc\u00de are rigidly connected to the cutter and generating gear, respectively, while the auxiliary coordinate systems Sa and Sb describe the relative position between the two"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000638_j.proeng.2015.08.069-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000638_j.proeng.2015.08.069-Figure5-1.png",
        "caption": "Fig. 5. The finite element meshing.",
        "texts": [
            " Stresses decrease then, as we move away from the contact zone, to lower values and almost vanishes as we get close to the lower contact surface (Fig. 4). The material is considered to behave everywhere as a purely elastic isotropic material. Fringe constant f=0.43N/mm, Young\u2019s modulus (E1=210000 MPa, E2=15.9 MPa) and Poisson\u2019s ratios (\u03bc1=0.3, \u03bc2=0.45) respectively for the sphere and the parallelepiped are introduced in the finite element program. The mesh is refined in the neighbourhood of the contact zone (Fig. 5) in order to achieve better approximation of stresses. The retardation angle is obtained with the equation N2 . The different values of can be determined with the following relation (equation 3): 22 4/2 xyyxfe (3) The different values of sin2\u03c6/2 which represent the simulated isochromatic fringes can then be easily calculated along the z axis and displayed (Fig. 6). We can see in the first slice at z=0, which corresponds to the direction of the applied load, a concentration of fringes in the neighbourhood of the contact zone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000840_00423114.2015.1023319-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000840_00423114.2015.1023319-Figure6-1.png",
        "caption": "Figure 6. Constraints (a) and the forces from the car body (b).",
        "texts": [
            " In the FEM, the stiffness at a certain node is a result of a superposition of the stiffness properties of the elements adjacent to it. This leads to artefacts on the contact stresses diagram and makes the mesh inapplicable for contact problems. To overcome this difficulty, several uniform rows of six-node prismatic elements were introduced on the surface layer (Figure 5(c) and 5(d)). The finite element model has about 64,000 degrees of freedom. Further analysis has shown that to be sufficient to obtain smooth distributions of the contact stresses. The constraints are imposed as shown in Figure 6(a). The scheme of the load application should provide a realistic stress\u2013strain state in all elements of the model. The load is applied as a system of concentrated vertical forces as shown in Figure 6(b). Two torques are resulted D ow nl oa de d by [ G az i U ni ve rs ity ] at 1 1: 00 3 0 A pr il 20 15 Vehicle System Dynamics 863 from the difference between the corresponding vertical forces. The values of the loads come from the dynamics simulation. 5.2. Multibody model of the 18\u2013100 bogie for dynamics simulation The multibody model of the 18\u2013100 bogie (Figure 7) was developed in the Universal Mechanism (UM) software package and described in detail.[3] UM is general-purpose software for multibody system dynamics analysis often used for the simulation of railway vehicle dynamics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000822_j.proeng.2014.12.047-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000822_j.proeng.2014.12.047-Figure2-1.png",
        "caption": "Fig. 2. Tensegrity plate.",
        "texts": [
            " Struts were made of steel tubes with parameters: Es=210 GPa, As = 7.26 cm2, and cables of steel ropes with parameters: Es = 210 GPa, As = 2.01 cm2. The analysed module has one infinitesimal mechanism and a corresponding self-stress state, which stiffens the module. Cable removal causes the module to collapse. The further study considers a sample plate structure constructed from reversed 3-strut Simplex modules (described above), located in such a way that the left module is always joined with the right and vice versa (Fig. 2). No additional cables were used in this structure. The plate is supported in four nodes. It has one infinitesimal mechanism (Fig. 3), and self-stress states in each module [2]. Calculations of such structures cannot be performed using geometrically linear theory because of the infinitesimal mechanism. Finite element analysis according to the second order theory allows to consider prestressing forces (in one module or several) through geometric stiffness matrix, and to obtain structural response for external loading"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002948_s40435-020-00624-z-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002948_s40435-020-00624-z-Figure2-1.png",
        "caption": "Fig. 2 Suspended cable robot with reconfigurable pulley mechanism",
        "texts": [
            " The cable robot is able to reconfigure its contact points thanks to the pulley mechanism as shown in Fig. 1. Two robot structures are selected based on the optimization results in [20,21]. They will be compared during the reconfiguration to determine theminimumcable tensions and velocities. This paper is organized as follows: the geometric model of the cable robot is presented in Sect. 2. Then, the kinematic and dynamic analyses are carried out in Sects.3 and 4 whose results will be used for trajectory generation and reconfiguration planning in Sects.5 and 6. Figure 2 shows the geometric model of a suspended cable robot with pulley mechanism. It consists of a base frame and a moving platform that are connected by four cables with reconfigurable pulleys. The base frame is a cube of length lb, widthwb and height hb. At the bottomof it, a fixed coordinate is located which is denoted by Fb with axes x, y, z. The moving platform is a cube of length l p, width wp and height h p. On top of it, a local coordinate is placed which is denoted by Fp of axes u, v, w. The geometric center of the moving platform is denoted by point P and its position coordinate expressed in the base frame is bp = [ x y z ]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001504_amm.459.449-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001504_amm.459.449-Figure1-1.png",
        "caption": "Fig. 1 Single Link Solid Soft Finger with Hemispherical Fingertip",
        "texts": [
            " Modal analysis has become a wide spread means of finding modes of vibration of a structure. Hence, finite element (FE) modal analysis was conducted on the finger model from which the first ten fundamental frequencies and mode shapes were computed. Dynamic characteristics such as amplitude, velocity and acceleration with respect to time in the selected excitation frequency range were then computed from transient analysis. The soft finger was modeled as a solid cylindrical single link with hemispherical fingertip. The model was initially done in Solid works \u00ae 2010 as shown in Fig.1 and was then exported to ANSYS \u00ae V12 for FE analysis. SILASTIC P1, Silicone RTV was the material used in the analysis. It is an elastomer in the form of liquid, when mixed with curing agent and allowed to cool at room temperature, it will become a solid rubber for the preparation of products. The stress-strain behaviour of hyper elastic elastomeric materials is nonlinear in nature and hence the linear elastic modulus cannot be applied in static and dynamic analysis. Nonlinear material constants are to be computed from any one of the experimental studies such as uniaxial compression, uniaxial tensile test, biaxial test, shear test etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003893_b978-0-12-803224-4.00289-2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003893_b978-0-12-803224-4.00289-2-Figure3-1.png",
        "caption": "Figure 3 Sandwich cell for transmission measurement under semi-infinite linear diffusion conditions. Reprinted by courtesy of Marcel-Dekker, Inc. from Kuwana T and Winograd N (1974) Spectroelectrochemistry at optically transparent electrodes. I. Electrodes under semi-infinite diffusion conditions. In: Bard AJ (ed.) Electroanalytical Chemistry. A Series of Advances, vol. 7, pp. 1\u201378.",
        "texts": [
            " In the semi-infinite linear diffusion spectroelectrochemical cell geometry (Figure 1a), the cell is analogous to a conventional electrochemical cell, the electrode being in contact with solution much thicker than the diffusion layer adjacent to the electrode. Semi-infinite linear diffusion spectroelectrochemical cell design requires that electrolysis products generated at the counterelectrode should not interfere with the absorbance measurement and that complete deoxygenation should be easily achieved. Figure 3 shows the classic sandwich-cell design, with a thin film OTE as the working electrode. The reference electrode and counterelectrode and the side arms for degassing are positioned so that the cell may be placed with the surface of the OTE in a horizontal plane. A Luggin capillary places the reference electrode near to the surface of the OTE for minimization of solution resistance in the control of the working electrode potential. These cells are normally used for experiments (chronoamperometry and chronocoulometry) in which large-amplitude steps are applied in order to carry out an electrolysis in the diffusion region"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001798_0954406216631370-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001798_0954406216631370-Figure1-1.png",
        "caption": "Figure 1. Sub-system of gear system.",
        "texts": [
            " This drag torque on the idler gear is still poorly understood and an analytical formulation of the problem could be useful in order to evaluate its influence. The purpose of the present work is to investigate the dynamical behavior at different drag torque by considering the double-sided lubricant films. All these will increase our understanding of its roles during gear engagement, and help to identify improvements that can be made to the mechanism for the sake of reducing the vibration and noise of the gear pairs system. The system in this study consisted of two gears mounted on well-aligned input and output shafts as shown in Figure 1. The gears were standard errorless involute helical gears with no modifications, and the gear pair was modeled as a generalized lumped parameter torsional vibration system. In theory, this concept is also applicable to helical, spiral bevel, and worm gears as long as the geometrical configuration is represented appropriately. In the initial arrangement, the pinion tooth is placed in a central relative position, where the tooth separation is half of the backlash on both side. And the centers of both gears are not allowed to move laterally"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure2-1.png",
        "caption": "Fig. 2 Physical model of driving system on servo press.",
        "texts": [
            " In order to solve the nonlinearities problem, the dynamical model of servo press is derived using computed torque module and PD parameters adaptively based on the change of angle error, speed error and reference acceleration in the second part. Thirdly, the PTC controller is designed for this model which is disturbed by unknown mathematical characteristics. Finally, through simulations, the proposed scheme shows valid control performance. II. DYNAMIC MODEL OF SERVO PRESS A press mechanism driven by an IM is shown in Fig.1 and Fig.2, where m1, m2, m3, mc, ms2 and ms are the masses of the three connecting rods, the crank, shaft2 and the slide, respectively. L1, L2 and L3 are the lengths of connecting rods, respectively. R is the radius of rotating crank. d and h are the horizontal and vertical distances from O to B. The slide can move along the guides up and down when the IM shaft rotates. The main task of this research is to track the desired motion of This study is supported by the \u201cKey Innovation Special Program\u201d of Qilu University of Technology"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002745_s12206-019-1235-8-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002745_s12206-019-1235-8-Figure7-1.png",
        "caption": "Fig. 7. L-shaped path to be followed by the moving platform.",
        "texts": [
            " The platform moves in three directions of x, y and z, whereas it does not rotate about any of those axes. All the parameters are displayed in the Table 1 above. The initial conditions taken here for the three dimensional hybrid manipulator are shown in Table 2. All the values in z direction are taken zero here because the path which we will consider later to be followed by the manipulator is a two-dimensional L-shaped path in x-y plane and force control will also be done for this trajectory. The path to be followed by the hybrid manipulator is shown in Fig. 7 where point G of the upper movable platform will trace the path from A to C. The force or torque interaction between the tool attached on the platform and the environment is measured by the sensor measuring force or torque that is fixed on the upper movable platform of the hybrid manipulator. This force of interaction is compared with the restricted force value given in Table 1 for force controller and then the compensation gain z is used to adjust the impedance at the point of interaction. The simulation is done for 4 seconds",
            " time at the time of controlling force without removal of amnesia are displayed in Figs. 8(a) and (b), respectively. The force control and amnesia error is displayed in these figures. In both these cases, the tool is not able to come at the starting point after finishing one cycle. It is clearly seen from Fig. 8 that initially, the reference trajectory given to the manipulator is exactly tracked in x-direction as shown in Fig. 8(a), i.e. the point G of the platform moves from point A to B as shown in Fig. 7. During this time, force control is done in y-direction. Also, there is some difference in the command and response of y-displacement for this time duration, which is called amnesia error as shown in Fig. 8(b). Then the platform moves from point B to point C as shown in Fig. 5, i.e. in inertial Y-direction so that trajectory tracking is done fully in direction of y-axis as displayed in Fig. 8(b). The force control is done in x-direction for this time after which there comes some difference in command and response of x-displacement, known as amnesia error as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002933_0954406220916536-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002933_0954406220916536-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of cotton-picker HMCVT. Note: nin, nout are the speeds of input and output shaft, respectively; i1\u2013i3 are transmission ratio of each gear pair; K1, K2 represent planet gears; C1, C2 represent shafts of clutches; B represents shaft of brake? P, M represent variable displacement pump and fixed displacement motor, respectively.",
        "texts": [
            " A speed torque compensation control was used to solve the problem of output speed fluctuation caused by load change, which provides a reference for realizing the quick and smooth change of the HMCVT speed ratio of the cotton picker. A self-developed HMCVT was used in self-propelled cotton pickers. Depending on the complex environment in which cotton is picked, the constant changes in speed and load were required during the operation. A new HMCVT was developed that combined a volume speed control loop of \u2018\u2018variable displacement pump\u00fe quantitative motor\u2019\u2019 and a Simpson planetary gear train, as shown in Figure 1. Characteristics of continuous section change of transmission When the vehicle starts, the HMCVT is in the H section and the brake B locks the K2 planetary gear ring. At this time, the power is transmitted to the sun gear in the planetary row K2 through the variable pump\u2013 motor system, and the power is output through the planetary carrier in the planetary row K2. Currently, the output speed of the HMCVT is noH. When the speed exceeds the maximum speed of the pure hydraulic section, the HMCVT switches to the HM1 section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003484_0954406220963148-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003484_0954406220963148-Figure7-1.png",
        "caption": "Figure 7. Schematic diagram of path deflection angle.",
        "texts": [
            " The schematic diagram is shown in Figure 6, where dh can be calculated as: dh \u00bc Aixh \u00fe Biyh \u00fe Ciffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Ai \u00fe Bi p (15) where dh is the distance from the path to the h-th obstacle; (xh, yh) is the coordinate of the the h-th obstacle\u2019s circle center; Ai, Bi, Ci are the coefficients of the path between the i-th point and the (i 1) point of the bat individual. If the requested path does not pass an obstacle, K\u00bc 0, otherwise, K\u00bc 100. K\u00bc 100 dh < rh 0 dh > rh (16) where rh is the radius of the h-th round obstacle The formula of the penalty term in the fitness function is as follows: Lpunish\u00bc XD i\u00bc1 K (17) In order to improve the smoothness of the planned path, reduce the turning points and avoid fluctuations and shocks caused by the robot\u2019s excessive turning angle during operation, the deflection angle at each node (as shown in Figure 7) is calculated to measure the path\u2019s smoothness. Langle \u00bc XD i\u00bc1 180 arccos li ! li\u00fe1 ! li ! li\u00fe1 ! (18) where Langle represents the sum of the path deflection angles;~li represents the vector formed by the i-th and i 1 points of the bat individual; ~li\u00fe1 represents the vector formed by the i\u00fe 1th and i-th points of the bat individual; ~li and ~li\u00fe1 are the modulus of two vectors, that is, the length of two adjacent line segments at the i-th point. In summary, the fitness function constructed in this paper is expressed as: Fitness\u00f0path\u00de \u00bc f1Llength \u00fe f2Lpunish \u00fe f3Langle (19) where f1, f2, f3 represent the weight coefficients in the path length, obstacle penalty term, and path smoothness function, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003021_b978-0-12-418691-0.00003-4-Figure3.15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003021_b978-0-12-418691-0.00003-4-Figure3.15-1.png",
        "caption": "FIGURE 3.15 Beam undergoing flexural vibration: (a) general layout; (b) free-body diagram of an infinitesimal beam element; (c) stress distribution across the thickness; (d) horizontal displacement induced by flexure.",
        "texts": [
            " At the point of maximum displacement, the elastic energy reaches a maximum while the kinetic energy is zero, whereas at the point of zero displacement (but maximum velocity), the elastic energy is zero while the kinetic energy reaches a maximum. Recall that the assumption of natural modes of vibrations is that the complete body undergoes the vibration in phase, i.e., all the points in the body reach the maximum amplitude and then the zero amplitude, etc., in the same time. 3.4 FLEXURAL VIBRATION OF A BEAM Consider a uniform beam of length l, mass per unit length m, bending stiffness EI, undergoing flexural vibration of displacement w\u00f0x, t\u00de as shown in Figure 3.15a. Centroidal axes are assumed. An infinitesimal beam element of length dx is subjected to the action of STRUCTURAL HEALTH MONITORING WITH PIEZOELECTRIC WAFER ACTIVE SENSORS bending moments, M\u00f0x, t\u00de, M\u00f0x, t\u00de1M0\u00f0x, t\u00dedx, axial forces, N\u00f0x, t\u00de, N\u00f0x, t\u00de1N0\u00f0x, t\u00dedx, and shear forces V\u00f0x, t\u00de, V\u00f0x, t\u00de1V0\u00f0x, t\u00dedx (Figure 3.15b). We will briefly review the Euler-Bernoulli theory of bending. (Shear deformation and rotary inertia effects are ignored.) Free-body analysis of the infinitesimal element of Figure 3.15b yields N0\u00f0x, t\u00de5 0 (311) V0\u00f0x, t\u00de5m \u20acw\u00f0x, t\u00de (312) M0\u00f0x, t\u00de1V\u00f0x, t\u00de5 0 (313) Differentiation of Eq. (313) and substitution into Eq. (314) yields, upon rearrangement, M00\u00f0x, t\u00de1m \u20acw\u00f0x, t\u00de5 0 (314) The N and M stress resultants are evaluated by integration of the direct stress across the cross-sectional area shown in Figure 3.15c, i.e., N\u00f0x, t\u00de5 \u00f0 A \u03c3\u00f0x, z, t\u00dedA (315) M\u00f0x, t\u00de52 \u00f0 A \u03c3\u00f0x, z, t\u00dezdA (316) Using the stress-strain constitutive relation of Eq. (164), the axial force and moment stress resultants (315) and (316) can be expressed as N\u00f0x, t\u00de5E \u00f0 A \u03b5\u00f0x, z, t\u00dedA (317) M\u00f0x, t\u00de52E \u00f0 A \u03b5\u00f0x, z, t\u00dezdA (318) Kinematic analysis (Figure 3.15d) yields the direct strain \u03b5\u00f0x, z, t\u00de in terms of the flexural motion w\u00f0x, t\u00de and the thickness-wise location z, i.e., u\u00f0x, z, t\u00de52zw0\u00f0x, t\u00de \u03b5\u00f0x, z, t\u00de5 u0\u00f0x, z, t\u00de52zw00\u00f0x, t\u00de (319) Substitution of Eq. (319) into Eqs. (317), (318) and integration over the area yields N\u00f0x, t\u00de52Ew00\u00f0x, t\u00de \u00f0 A zdA5 0 (320) M\u00f0x, t\u00de5Ew00\u00f0x, t\u00de \u00f0 A z2dA (321) STRUCTURAL HEALTH MONITORING WITH PIEZOELECTRIC WAFER ACTIVE SENSORS Note that Eq. (320) indicates that the axial stress resultant is zero, i.e., N\u00f0x, t\u00de5 0, since centroidal axes were assumed",
            " (428), (429), (432), it seems appropriate to express Eq. (426) as Wj\u00f0x\u00de5Aj\u00bd\u00f0cosh\u03b3jx2 cos\u03b3jx\u00de2\u03b2j\u00f0sinh\u03b3x2 sin\u03b3x\u00de Wj\u00f0x\u00de5Aj\u00bd\u00f0cosh\u03b3jx2 cos\u03b3jx\u00de2\u03b2j\u00f0sinh\u03b3jx2 sin\u03b3jx\u00de 5Aj\u00bd2cos\u03b3jx1 \u03b2jsin\u03b3jx1 1 2 \u00f012\u03b2j\u00dee\u03b3jx 1 1 2\u00f011 \u03b2j\u00dee2\u03b3jx (433) Assume a uniform beam undergoing flexural vibration under the excitation of a timedependent distributed transverse excitation f\u00f0x, t\u00de, as shown in Figure 3.18. The units of f\u00f0x, t\u00de are force per length. Free-body analysis of an infinitesimal element similar to that presented in Figure 3.15b yields V0\u00f0x, t\u00de1 f\u00f0x, t\u00de5m \u20acw\u00f0x, t\u00de (434) M0\u00f0x, t\u00de1V\u00f0x, t\u00de5 0 (435) STRUCTURAL HEALTH MONITORING WITH PIEZOELECTRIC WAFER ACTIVE SENSORS STRUCTURAL HEALTH MONITORING WITH PIEZOELECTRIC WAFER ACTIVE SENSORS Differentiation of Eq. (435) and substitution into Eq. (434) yields, upon rearrangement, M00\u00f0x, t\u00de1m \u20acw\u00f0x, t\u00de5 f\u00f0x, t\u00de (436) Substitution of Eq. (322) into Eq. (436) yields the equation of motion for forced flexural vibration of a beam under transverse excitation, i.e., EI w0000\u00f0x, t\u00de1m \u20acw\u00f0x, t\u00de5 f\u00f0x, t\u00de (437) Without loss of generality, we assume the external excitation to be time harmonic in the form f\u00f0x, t\u00de5 f\u0302\u00f0x\u00deei\u03c9t (438) and obtain m \u20acw\u00f0x, t\u00de1EI w0000\u00f0x, t\u00de5 f\u0302\u00f0x\u00deei\u03c9t (439) Assume the vibration response w\u00f0x, t\u00de to be expressed as a series expansion in terms of vibration modes Wj\u00f0x\u00de that satisfy the free-vibration equation of motion Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002744_tmag.2019.2950880-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002744_tmag.2019.2950880-Figure7-1.png",
        "caption": "Fig. 7. Thermal LPM of the magnetic circuit.",
        "texts": [
            " (2) The new B\u2013H effective (B\u2013Heff) curve can be computed by assuring the same density energy, in time and in frequency domains. With this, knowing that the magnetic-flux density is imposed with a sinusoidal curve with the effective value of Bef , the equivalent magnetic-field value, H\u0302eff , and its equivalent magnetic permeability, \u03bc\u0302eff , can be computed as follows: H\u0302eff = 1 Bef 8 T \u222b T/4 0 (\u222b B(t) B(0) H (B)d B ) dt (3) \u03bc\u0302eff = B2 ef 8 T \u222b T/4 0 (\u222b B(t) B(0) H (B)d B ) dt . (4) The magnetic circuit\u2019s thermal model was built based on an LPM (see Fig. 7), where Plat , Ptop, and Pfr are the power losses dissipated through lateral, top, and frontal surfaces with thermal convective resistances Rconvlat , Rconvtop , and Rconvfr , respectively. This allows a good approximation of the magnetic circuit\u2019s steady-state temperature, with fast computation times. Due to the circuit\u2019s main composition (with metallic materials), the conductive heat transfer in the magnetic core and copper windings can be neglected when compared with the convective heat transfer from its surface to the surrounding ambient"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003257_j.ijnonlinmec.2020.103550-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003257_j.ijnonlinmec.2020.103550-Figure1-1.png",
        "caption": "Fig. 1. Design of the system.",
        "texts": [
            " We note that this is the first known case where the Melnikov integral has been explicitly calculated for a nonholonomic system that has no invariant measure (i. e., for a nonconservative system). Previously [30] the Melnikov integral was calculated only for a formal nonholonomic system which preserves an invariant measure. https://doi.org/10.1016/j.ijnonlinmec.2020.103550 Received 22 February 2020; Received in revised form 3 June 2020; Accepted 10 July 2020 Available online 14 July 2020 0020-7462/\u00a9 2020 Elsevier Ltd. All rights reserved. Consider a system moving on a horizontal plane and consisting of two bodies (see Fig. 1): \u2014 a homogeneous spherical shell; \u2014 a frame which is a dynamically asymmetric rigid body fastened inside the shell. Let the frame be balanced relative to the shell, so that the center of mass of the entire system is at the geometric center of the shell. Also, assume that the frame relative to the shell rotates with constant angular velocity \u2126 about the axis passing through the geometric center of the shell \ud835\udc36. If the shell rolls without slipping, then in the moving coordinate system \ud835\udc36\ud835\udc651\ud835\udc652\ud835\udc653 (attached to the frame) the problem reduces to investigating an autonomous system of equations governing the evolution of the angular momentum \ud835\udc74 and the normal \ud835\udf38 at the contact point \ud835\udc43 (see [29] for details): "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003115_physrevfluids.5.053103-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003115_physrevfluids.5.053103-Figure9-1.png",
        "caption": "FIG. 9. Spatial validation. (a) Flow field generated by 60 Stokeslets (red arrows) shown as streamlines. (b) The absolute error between the exact solution and the numerical solution with a total of about 4000 Gaussian quadrature points. Color code represents log10(|uexa \u2212 unum|). (c) The l\u221e norm of the flow field shown as a function of the number of quadrature points.",
        "texts": [
            " (A3) It is easy to see that T\u0303i jk converges to the singular traction kernel Ti jk in the limit \u03b5 \u2192 0; the correction term induced by the regularization appears at O(\u03b52) and higher orders of \u03b5. To validate our boundary integral method, we construct a boundary value problem and test the algorithm against the exact solution. Specifically, we place 50 Stokeslets with random strengths inside the inner channel boundary and one Stokeslet with random strength at the center of each of the ten particles as shown in Fig. 9(a). The flow field uexa( ) created by these Stokeslets can be found by evaluating directly using the free-space Green\u2019s function. To obtain the numerical solution, we set the rigid-body velocity vector U\u2217 to be zero and treat the flow field on the channel walls and the particle surfaces, given by uexa(\u2202 ), as the boundary conditions on \u2202 where \u2202 \u2261 \u222a \u03b3 . Symbolically, one can think of uexa as uc and substitute it into (23) to solve for the corresponding density function \u03bc. The numerical solution unum( ) = u + u\u03b3 is then found by substituting \u03bc into (8) and (16). 053103-15 The logarithm of absolute error between uexa and unum is shown in Fig. 9(b) with about 4000 Gauss-Legendre quadrature points on and \u03b3 in total. It is noticeable that the algorithm has at least a 14-digit accuracy for most of the locations, and 12-digit accuracy is achieved even close to the particles. The l\u221e norm of the error as a function of the number of quadrature points is shown in Fig. 9(c). Next, we place N = 32 cilia with phase difference \u03c6 = 2\u03c0/N = \u03c0/16 as in (1) and use a standard Runge-Kutta fourth-order (RK4) scheme to march forward in time. Due to the lack of an exact solution in this case, we test the self-convergence rate with respect to t . We monitor the motion of a rigid particle of radius rp = 0.4 initially centered at (0,4.5) for a full cycle t \u2208 [0, 1] and for t = {0.04, 0.02, 0.01}. The particle is discretized using 128 quadrature points. At the final time T = 1, we measure the following quantities in Table II: Ex(T, t ) = \u2212 log2 |x t c (T ) \u2212 x t/2 c (T )|, Ey(T, t ) = \u2212 log2 |y t c (T ) \u2212 y t/2 c (T )|, E (T, t ) = \u2212 log2 | t (T ) \u2212 t/2(T )|, (B1) where (t ) = \u222b t 0 \u03c9 dt is the orientation of the particle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure3-1.png",
        "caption": "Fig. 3. Conventional C-shaped cores MB model.",
        "texts": [
            " However, in order to stably generate sufficient suspension force for suspending a shaft of a high-output motor, the general homopolar-type MB requires a large magnet and arch-shaped windings, as shown in Fig. 2(a). It is noted that the large magnet is required for generating the bias flux as the dc magnetic field and the arch-shaped windings is needed to achieve the high fill factor. Consequently, the general homopolar-type MB has a complicated structure and becomes costly. In order to realize the MB structure with both lower iron loss and lower cost than the general homopolar-type MB, our research group focuses on the homopolar-type MB structure using four C-shaped cores, as shown in Fig. 3 [2]\u2013[4]. The homopolar-type MB structure with the four C-shaped cores has potential for lower rotor iron loss than the general homopolar-type, because the magnetic field on a rotor core is dc, and the width between the magnetic poles can be smaller than that of the general homopolar-type, as shown in Fig. 3. However, if the homopolar-type MB using C-shaped cores is applied to a large-sized MB for high output, it is very complex to assemble four separated C-shaped cores. In addition, it is very difficult to increase the stator inner surface accuracy. 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In order to realize the stable magnetic suspension and the low loss, it is important that the stator inner accuracy be highly maintained",
            " This paper proposes a novel homopolar-type MB structure unifying the four C-shaped cores for high output, low rotor iron loss, and simple assembly. Fig. 4(a) shows the novel homopolar-type MB structure. The stator core of the novel homopolar-type is divided into stator bulk cores and stator laminated cores. The windings are wound to the stator bulk cores. The stator-laminated cores are made up of unifying teeth parts, in which the magnetic flux flows to the radial direction in the four separated C-shaped cores, as shown in Fig. 3. Therefore, assembling four separated C-shaped cores is not necessary. In addition, the width between the magnetic poles, shown in Fig. 4(b), easily becomes small by the unified stator laminated cores. It is noted that the leakage flux between the magnetic poles is blocked by inserting small magnets between the magnetic poles, as shown in Fig. 4(b). Since the novel homopolar-type MB has no large magnets for generating bias flux, the magnet volume compared with that in the general homopolar-type can be decreased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure1-1.png",
        "caption": "Fig. 1 The schematic diagram of gear pump",
        "texts": [
            " Third, the basic theory of blind source separation and particle swarm algorithm are studied, and the frequency measurement procedure of gear pump gear is designed. Finally, the simulation on multiple cracks of gear pump gear is carried out, and results show that the wavelet finite element can identify the location and size of the multiple cracks of gear pump gear correctly. Keywords Gear pump gear Wavelet finite element method Blind source separation Multiple cracks Introduction The gear pump is a common hydraulic pump, which can convert the mechanical energy into fluid pressure energy, and the schematic diagram of gear pump is shown in Fig. 1. A pair of gears is installed in the pump shell, which is covered by the end shields, and many sealed spaces between the pump shell, covers and teeth, slot are formed. When the drive shaft and reaction shaft rotate in the direction as shown in Fig. 1, the meshed gear tooth come away gradually, and the gear tooth exits from slot, and the volume of the sealed space will increase gradually to form partial vacuum, then the oil in the tank enters into sealed space from intake under action of atmospheric pressure, then the slot between two meshed teeth is filled. The oil is taken to the left with gear rotating. Because the gear teeth mesh gradually in the left sealed space, the sealed volume decreases accordingly, and the oil is extruded out of the seal space to discharge"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001316_066101-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001316_066101-Figure5-1.png",
        "caption": "Figure 5. Schematic diagram of the lattice structure.",
        "texts": [
            " There are N hexahedral units in the entire lattice structure, with O units along the X axis direction, P units along the Y axis direction, and Q units along the Z axis direction; thus N = O\u00d7 P\u00d7 Q. Vvirtual object = V1 + 2V2 + 2V3 + 2V4. (1) From the given conditions we obtain: V1 = (a\u2212 2h1)(b\u2212 2h2)(c\u2212 2h3). (2) As AB = a,BD = b, and using figure 4 we obtain: A\u0304B\u0304 = a\u2212 2h1 B\u0304D\u0304 = b\u2212 2h2 (3) and V2 = (a\u2212 2h1)(b\u2212 2h2)h3. (4) In the same way we obtain: V3 = (a\u2212 2h1)(c\u2212 2h3)h2, (5) V4 = (b\u2212 2h2)(c\u2212 2h3)h1. (6) As shown in figure 5, it is known that the lattice structure has N hexahedral units giving a volume NV1. In addition, as there are O hexahedral units in the X axis direction, P units in Laser Phys. 23 (2013) 066101 J Sun et al the Y axis direction and Q units in the Z axis direction, thus there are contributions of (Q + 1)OP V2; (O + 1)QP V3; and (P+ 1)OQ V4 and the total for all virtual objects is Vall virtual object = NV1 + (Q+ 1)OPV2 + (O+ 1)QPV3 + (P+ 1)OQV4. (7) The overall volume can be obtained from figure 5 as Vtotal volume = [(O+ 1)h1 + O(a\u2212 2h1)] \u00d7 [(P+ 1)h2 + P(a\u2212 2h2)] \u00d7 [(Q+ 1)h3 + Q(c\u2212 2h3)]. (8) The porosity is given by: \u03b7porosity = Vall virtual object Vtotal volume = OPQ(V1 + V2 + V3 + V4)+ OPV2 + QPV3 + OQV4 (Oa\u2212 Oh1 + h1)(Pa\u2212 Ph2 + h2)(Qc\u2212 Qh3 + h3) . (9) When Q = O = P = 3 \u221a N, h1 = h2 = h3, a = b = c, then the designed hexahedral unit is a regular hexahedral unit, and equation (9) can be simplified as: \u03b7porosity = ( 3 \u221a N)2(l\u2212 2h)2( 3 \u221a Nl+ 3 \u221a Nh+ 3h) ( 3 \u221a Nl\u2212 3 \u221a Nh+ h)3 . (10) 3.2.1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure17-1.png",
        "caption": "Fig. 17. The Singular configuration III.",
        "texts": [
            " 15 ), when the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R2 and J R6 are in the same direction, the joint axes of R joints J R4 and J R8 are in the same direction, but the two groups directions of joint axes are opposite during the movement. Then the mechanism makes a planar motion, which is called plane-motion Mode IV. Its typical configurations are shown in Fig. 16 . When the axes of R joints J R2 and J R8 coincide, the axes of R joints J R4 and J R6 coincide, and the axes of R joints J R3 and J R7 coincide, the configuration is a special configuration of Mode II. Or when the axes of R joints J R2 and J R4 coincide, the axes of R joints J R6 and J R8 coincide, and the axes of R joints J R1 and J R5 coincide, as shown in Fig. 17 , the configuration is also a special configuration of Mode II. Since the mechanism is an open-chain mechanism with 3 links and 2 R joints, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration III ( Fig. 17 ), when the axes of R joints J and J coincide, the axes of R joints J and J coincide, and the axes of R joints J and R2 R4 R6 R8 R1 J R5 coincide during the movement, the mechanism makes a planar motion, which is called plane-motion Mode V. Its typical configurations are shown in Fig. 18 . Due to the characteristics of antiparallelogram unit ( Fig. 19 ), it is a bifurcation point at \u03b8AA = 0 or \u03b8AA = \u03c0 . At a bifurcation point, an antiparallelogram could turn into a parallelogram. Therefore, there are some more special configurations as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003666_012019-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003666_012019-Figure4-1.png",
        "caption": "Figure 4.The maximum deformation of Tic-tac-toe type\u3001PSL-type and Optimized-type(from left to right).",
        "texts": [],
        "surrounding_texts": [
            "MCTE 2020\nJournal of Physics: Conference Series 1678 (2020) 012019\nIOP Publishing\ndoi:10.1088/1742-6596/1678/1/012019\n\u00a0 \u00a0 \u00a0 \u00a0 \u00a0 \u00a0\nextraction, based on the theory of tensor topology in the degradation area boundary and the dense regions of \"advantage principal stress line\" the two special cases, extraction principle is as follows:\nStep1. Import the principal stress line, and calculate every principal stress contour of maximum curvature kmax, will kmax between kmax + \u0394k together as one family with one assignment.\nStep2. Use function to solve the distance calculation on 3dmax every race of two principal stress boundary line of maximum point to point distance di and principal stress line number mi, linear density values for each group principal stress di/mi, curve was generated by size order PCi  {PC0, PC1, PCi}.\nStep3. Find out the maximum point-to-point distance between the initial principal stress line CURSi and the next principal stress line CURj of each group. Remove the redundant principal stress line if it does not meet the requirements, + is determined by the overall weight of the structure.\nIn this paper, three parameter values affecting the maximum deformation D are extracted and parameterized: shell thickness DS_T, stiffener width DS_W and stiffener height DS_H. In consideration of surface curvature and actual working conditions, a reasonable variable value range is set to continue the dimensional optimization analysis. After the optimal solution is obtained, the size is applied to the structural design. The influence of three parameters on the maximum deformation D is shown in Figure 3.\nmin D \uff081\uff09\nc ts t        \uff082\uff09\nD ( _ , _ , _ )f DS T DS W DS H \uff083\uff09",
            "MCTE 2020\nJournal of Physics: Conference Series 1678 (2020) 012019\nIOP Publishing\ndoi:10.1088/1742-6596/1678/1/012019\n\u00a0 \u00a0 \u00a0 \u00a0 \u00a0 \u00a0\nBy analysing the response surface between the variable and the response value, the influence rule of size parameters on the maximum deformation in the design of high rigid thin-wall structure is systematically analysed. These trends indicate that there is an interaction between the three variables, and the optimal solution exists for maintaining the mass within a certain range while obtaining the minimum deformation. The optimal solution: DS_W is 1.73mm, DS_H is 2.46mm, DS_T is 1.97mm, and the maximum deformation and mass are 0.08mm and 0.52kg respectively.\nFor the stiffness evaluation criteria of special-shaped shell structure, an index called efficiency of specific stiffness structure is introduced after considering the factors such as stiffness, material and structure quality comprehensively [9]. The higher this value is, the higher the stiffness of the structure will be. The concrete definition is shown in formula 4, where E represents young's modulus, represents the maximum deformation of the structure, and M represents the mass of the structure.\nE\nm   (4)\nIn the process of finite element analysis, the control group chose to use the traditional wellbore reinforcement model, while the experimental group chose to use the PSL reinforcement model along the main stress line and the optimized size model. The data in Section 2.3 was selected as the size of the optimization group. The weights of the three groups were kept basically the same. The same material was used, the same boundary conditions and loads were added, and the same grid was divided to carry out the comparative test.\nAmong them, the specific stiffness of PSL structure is 20% higher than that of the well-shaped structure, and the optimized structure is 40.9% higher than that of the traditional well-shaped structure. In addition, the total strain energy of the optimized structure [12] decreases to different degrees compared with the other two structures, which proves the feasibility of the method from the side.",
            "MCTE 2020\nJournal of Physics: Conference Series 1678 (2020) 012019\nIOP Publishing\ndoi:10.1088/1742-6596/1678/1/012019\n\u00a0 \u00a0 \u00a0 \u00a0 \u00a0 \u00a0\nThe three models were printed by the light curing technology, (a) adding traditional Tic-tac-toe type reinforcing rib to the model, (b) adding master stress line reinforcing rib to the model, and (c) adding optimized master stress line reinforcing rib to the model. Label the three models and carry out compression tests, as shown in figure 6. Due to the outer surface of the workpiece add aerodynamic load for special-shaped structure, by means of mechanical pressure on aerodynamic load is very difficult, special-shaped structure add and add the uniformly distributed load by means of hydraulic or pneumatic, need the workpiece sealing up and down, change the characteristics of the special-shaped structure itself, lost the significance of research, so we use the equivalent of the simplified load.\nThe relevant data after the workpiece is compressed is shown in figure 7, where the Tic-tac-toe type shape structure of the reference group appears obvious unloading phenomenon when the maximum load is 800N and the displacement is 8.5mm, and the compression stiffness [13] is 94N/mm. When the maximum load of PSL model is 870N, the displacement is 7.7mm, the maximum load capacity is 870N, and the compression stiffness is 112N/mm. When the maximum load of the optimization model is 910N, the displacement is 7.6mm, the maximum load capacity is 900N, and the compression stiffness is 120N/mm."
        ]
    },
    {
        "image_filename": "designv11_30_0001572_s11071-013-1014-5-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001572_s11071-013-1014-5-Figure2-1.png",
        "caption": "Fig. 2 Computation model of MQL deep hole drilling: (a) computation model of MQL deep hole drilling in xoy plane; (b) coordinate of drilling shaft",
        "texts": [
            " Based on a similar idea, to solve the hydrodynamic pressure function of cutting fluid in MQL deep hole drilling, the authors consider the compressible fluid Reynolds equation in oil/air feature arising from cutting fluid to satisfy the finite length drilling shaft 1 r2 \u2202 \u2202\u03b8 ( p\u03b43 \u03bc \u2202p \u2202\u03b8 ) + \u2202 \u2202z ( p\u03b43 \u03bc \u2202p \u2202z ) = 6\u03c9 \u2202(p\u03b4) \u2202\u03b8 + 12 \u2202(p\u03b4) \u2202t (1) where r is the drilling shaft radius; \u03b8 is the circumferential coordinate of drilling shaft; \u03bc is the dynamic viscosity of cutting fluid; \u03b4 is the thickness of air/oil mist, \u03b4 = c(1 + \u03b5 cos\u03d5); \u03b5 is the transverse displacement ratio of drilling shaft; c is radial clearance between the drilling shaft and deep hole surface; \u03d5 is the angle of drilling shaft whirling; p is the hydrodynamic pressure function of cutting fluid; t is time; \u03c9 is the rotational speed of drilling shaft. The computation model of MQL deep hole drilling is depicted in Fig. 2. On the basis of the assumption that the flow is isothermal and laminar, the boundary conditions of MQL fluid are stated as follows: (1) The pressure p inside deep hole on inlet section and outlet section equals the supply pressure (pin); (2) The pressure p is a continuous and symmetrical function at the half of drilling depth (l/2). (3) The pressure p is a periodic function for \u03b8 . In order to facilitate the calculation, we define the following dimensionless variables: z = z\u0304l/2, \u03b4 = \u03b4\u0304c = c(1 + \u03b5 cos\u03d5), p = p\u0304pa, t = \u03c9t\u0304 (2) where pa is the atmospheric pressure, l is the drilling depth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003481_icra40945.2020.9197323-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003481_icra40945.2020.9197323-Figure4-1.png",
        "caption": "Fig. 4. Image showing the racetrack-shaped trajectory used to study the steering abilities of the swimmer (swimmer not to scale).",
        "texts": [
            " The sign of this variable determines the direction followed by the swimmer along the path. Abruptly changing this variable from a positive to a negative value causes the swimmer to quickly change direction (see Subsections IV-A and IV-C). This section experimentally measures the position error while steering as a function of turning radius, and measures the step-out frequency as a function of turning radius. For this experiment, the swimmer was programmed to navigate a racetrack-shaped trajectory placed in the horizontal plane at z = 0 (see Fig. 4). The swimmer was rotating at a constant speed along the trajectory. The navigation was performed for three turn radii and for eight rotational speeds. The positioning error is calculated using (2), where x, y and z are the coordinates of the position of the swimmer and xc, yc and zc are the coordinates of the closest point of the trajectory. E = \u221a (x\u2212 xc)2 + (y \u2212 yc)2 + (z \u2212 zc)2 (2) 7640 Authorized licensed use limited to: University of Canberra. Downloaded on October 04,2020 at 15:15:50 UTC from IEEE Xplore",
            " Restrictions apply. of the rotational speed of the magnetic field plane to a constant value, set to 15,000 rad/s2 in our experiment. The rotational direction of the swimmer must be reversed to keep the swimmer moving in the direction requested by the TP. For this reason, the orientation change request is also sent to the module calculating the rotational speed of the swimmer, which takes care of inverting the rotational direction. The swimmer was programmed to follow a racetrackshaped trajectory (see Fig. 4). The orientation was manually reversed using a toggle switch in the LabVIEW graphical interface. A video of this experiment is attached to this paper. Frames from the video as well as a plot of the experimental trajectory are presented in Fig. 10. For this maneuver the TP requests both an orientation and a direction reversal. In this case the orientation is flipped but the rotational direction of the swimmer does not change. This produced a 180\u00b0 change in the heading of the swimmer. The swimmer was programmed to follow a straight horizontal line and perform a combined direction and orientation reversal maneuver at each end until the program is manually stopped"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001515_acc.2014.6859324-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001515_acc.2014.6859324-Figure1-1.png",
        "caption": "Fig. 1: Top down view of a tailsitter pitching down 20\u25e6 and changing its heading by 180\u25e6",
        "texts": [
            " While the QF attitude error method has been used successfully in satellite attitude control [8], there are issues that arise when applying it to aircraft controllers that most designers never take into account. The key problem can clearly be seen in an example. Let the cur- rent attitude be vertical, \u03b7 = [ 1\u221a 2 0 1\u221a 2 0 ]T , and \u03b7d = Euler2Quat([0 70\u25e6 180\u25e6]), where Euler2Quat([\u03c6 \u03b8 \u03c8]) converts a ZYX Euler angle sequence to a quaternion. Ideally, in this situation, the attitude error would lie only in the body x-y plane which would cause a rotation similar to the one in Figure 1a. However, the QF error is e\u0303QF = [\u22120.9848 0 0.1736] T which lies in the body x-z plane. If the attitude controller used this error directly, for example in a PID loop, the aircraft would initially rotate towards the y axis, as shown in Figure 1b which would cause some undesired motion. This point is further illustrated in Figure 2 which shows e\u0303QF for \u03b7 = [ 1\u221a 2 0 1\u221a 2 0 ]T and \u03b7d = Euler2Quat([0 70\u25e6 \u03c8]) where \u03c8 \u2208 [\u2212180\u25e6 180\u25e6]. As can be seen, using e\u0303QF will cause the aircraft to have some undesired motion if the aircraft wants to pitch and change its heading at the same time. This problem occurs anytime the aircraft wants to rotate about a vector in the body y-z plane, which we will define as tilt, and change its heading. The RTT method was developed by Matsumoto et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-FigureA.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-FigureA.1-1.png",
        "caption": "Fig. A.1. The coordinate systems of two adjacent links connected by R joints.",
        "texts": [
            "1 The loop closure equation of an 8R mechanism The loop closure equation of the spatial single-loop 8R mechanism is as follows [15] : T 12 T 23 \u00b7 \u00b7 \u00b7 T 81 = I (A.1.1) where T i ( i + 1) is the transformation matrix from link i to link i + 1; when i = 8, i + 1 = 1; I is the identity matrix. The transformation matrix is as follows T (i +1) i = T \u22121 i (i +1) = \u23a1 \u23a2 \u23a3 cos \u03b8i sin \u03b8i 0 \u2212a i (i +1) \u2212 cos \u03b1i (i +1) sin \u03b8i cos \u03b1i (i +1) cos \u03b8i sin \u03b1i (i +1) R i sin \u03b8i sin \u03b1i (i +1) sin \u03b8i \u2212 sin \u03b1i (i +1) cos \u03b8i cos \u03b1i (i +1) \u2212R i cos \u03b8i 0 0 0 1 \u23a4 \u23a5 \u23a6 (A.1.2) The coordinate systems of two adjacent links connected by R joints are shown in Fig. A.1 : z i -axis is along the axis of R joint i , and x i -axis is along the common normal between axes z i \u22121 and z i . The D-H parameters are defined as follows: The length of link i ( i + 1), a i ( i + 1) , is the common normal distance from z i to z i + 1 along x i + 1 ; T he twist of link i ( i + 1), \u03b1i ( i + 1) , is the rotation angle from z i to z i + 1 about x i + 1 ; T he offset of joint i, R i , is the common normal distance from x i to x i + 1 along z i ; The joint variable of R joint i, \u03b8 i , is the rotation angle from x i to x i + 1 about z i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002540_ssrr.2016.7784286-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002540_ssrr.2016.7784286-Figure7-1.png",
        "caption": "Fig. 7. Target angle of subtracks in pitch down mode",
        "texts": [
            " In case the front subtracks contact the upper floor with the robot\u2019s position estimation error, it may hit the peak of the final step or the upper floor strongly. As a result, \u03b8p increases, and in the worst case, the falling backward mode occurs. To avoid the above situation, in this paper, we propose a method that absorbs the position estimation error for practical use. Specifically, the robot operates in a motion such that its height of free fall does not exceed the maximum acceptable height hmax for the robot and the surrounding environment when the robot touches down on the upper floor. Fig. 7-(A) indicates a state where the joint angles of the rear subtracks are equal to zero, and the zero moment point matches the peak of the final step. This is the state transition configuration for which the robot starts to fall down forward, and h indicates the height of free fall. Based on the geometric condition of Fig. 7-(A), the reference joint angle of the front subtrack should be \u03b8reff , described in Eq. (9) below, to make the height of free fall equal to the intended h less than hmax. \u03b8reff = arcsin ( lm 2 sin \u03b8s\u2212r(1\u2212cos \u03b8s)\u2212h\u2212d0 sin \u03b8s lf ) \u2212\u03b8s, (9) where d0 is the offset distance in the front-back direction of the robot between the peak of the final step and the center of the main tracks, as shown in Fig. 7-(A). On the other hand, the joints of the rear subtracks should be controlled downward slightly, to support the robot body falling forward smoothly, as shown in Fig. 7-(B). A suitable angle of the rear subtrack can be determined by the configuration of the robot in the figure: the three contact points do not generate the internal force to grasp the stairs. Depending on the geometric condition, the reference joint angle of the rear subtrack \u03b8refr can be calculated by : \u03b8refr = arctan ( p sin \u03b8s \u2212 (lm/2 + d0) sin \u03b8 0 p p cos \u03b8s \u2212 (lm/2 + d0) cos \u03b80p ) \u2212\u03b80p , (10) 978-1-5090-4349-1/16/$31.00 \u00a92016 IEEE 115 where \u03b80p is the pitch angle of the robot body when the front subtrack makes contact with the upper floor after the robot falls down forward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001143_acc.2013.6579933-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001143_acc.2013.6579933-Figure2-1.png",
        "caption": "Figure 2. Blade Flapping Motion",
        "texts": [],
        "surrounding_texts": [
            "Modeling is based on two key ideas. First, physics principles and appropriate assumptions are used to directly result in nonlinear ordinary differential equations (ODEs), without resorting to partial differential equations models. Second, flight and blade dynamics modes that are critical for safe and performant helicopter operation are captured. Multibody dynamics was used to include all helicopter components: fuselage, articulated main rotor with 4 blades, empennage, landing gear, tail rotor [4]. The key modeling steps are described next. 978-1-4799-0178-4/$31.00 \u00a92013 AACC 794"
        ]
    },
    {
        "image_filename": "designv11_30_0002850_s00170-020-05095-2-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002850_s00170-020-05095-2-Figure1-1.png",
        "caption": "Fig. 1 Structure of the spindle",
        "texts": [
            " However, in the existed research, few studies were related to these issues. In this study, an optimization method for heat transfer and thermal deformation is proposed to reduce the thermal elongation of high-speed spindle in a gear form grinding machine. The results of finite element analysis and experimental show that the heats transferred to the components that affect the accuracy of the spindle are very small, and the thermal elongation of the grinding wheel shaft is about zero. As shown in Fig. 1, the high-speed spindle of the gear form grinding machine is composed of three sections: one grinding wheel shaft B and two support shafts A and C. The grinding wheel shaft B is connected with the support shafts A and C through two diaphragm couplings, and the support shafts A and C are supported by three sets of angular contact ball bearings. The positional accuracy of the grinding wheel shaft affects the machining accuracy of machined gears, and the thermal deformations of the spindle components influence the positional accuracy of grinding wheel shaft",
            " Comparing of formulas 1 and 2, when the structure of the spindle was determined, the total thermal elongation of grinding wheel shaft depends on the average temperature rise of support shaft A, grinding wheel shaft B, and right diaphragm coupling. According to the heat transfer process of the spindle, the total thermal elongation of grinding wheel shaft when the left bearings are fixed is greater than that when the middle bearings are fixed. Therefore, the optimum constrained mode is that the middle bearings are fixed but the other two sets of bearings can move freely along the axial direction as shown in Fig. 1. It can be seen from formula 2 that reducing the heat transferred to the right diaphragm coupling and grinding wheel shaft can greatly reduce the thermal elongation of grinding wheel shaft. To obtain the minimum thermal elongation, the heat transfer of the spindle should be optimized. In order to reduce the heat transferred to the grinding wheel shaft and its associated components, the heat transfer in each component of the spindle should be analyzed firstly. Take the middle bearings as an example, the heats generated by middle bearings\u2019 rotation transfer to the sleeve firstly, then to the diaphragm coupling, and finally to the grinding wheel shaft as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001756_978-3-319-24502-7_10-Figure10.24-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001756_978-3-319-24502-7_10-Figure10.24-1.png",
        "caption": "Fig. 10.24 Schematics of an EWDLS set-up. The wave vectors of the incident evanescent and scattered light beams are ke and ks, respectively, with their difference defining the scattering vector q = ks\u2013ke. Independent experimental variation of the components qk and q? of the scattering vector parallel and perpendicular to the confining glass wall allows for the determination of wall-distance averaged diffusion coefficients. See the text for details. Redrawn after [95]",
        "texts": [
            " In many experiments such as in EWDLS, the relation D \u00bc kBTl; \u00f010:115\u00de between the hydrodynamic mobility matrix and the diffusion matrix, D, of a dispersed Brownian particle at system temperature T is used, on measuring the diffusion matrix coefficients instead of the associated hydrodynamic mobility coefficients. Light scattering is a powerful tool to investigate the properties of sub-micron soft matter systems [97]. In evanescent wave scattering experiments, a colloidal suspension is typically illuminated by a monochromatic laser beam that is totally reflected from a planar glass surface bounding the sample, so that no refracted light enters into the suspension (which is placed above the glass surface in Fig. 10.24) except for an evanescent wave whose intensity decays exponentially in going away from the glass surface into the suspension. Thus, only a colloidal particle close to the glass surface scatters enough of the incident evanescent light to be detected. The penetration depth, 2=j, of the evanescent wave can be changed to probe the particle diffusion at different glass surface-particle distances. For a review of the EWDLS method including experimental details, see [98]. The key quantity determined in (EW)DLS experiments is the scattered light intensity time-autocorrelation function",
            " This quantity can be theoretically predicted on basis of the generalized Smoluchowski equation determining the evolution of the configurational probability density function of Brownian particles under Stokes flow conditions [1, 99]. Inside the bulk region of a very dilute suspension of colloidal hard spheres far away from confining walls, the first cumulant is proportional to C \u00bc q2D0, where q is the modulus of the scattering vector q, and D0 \u00bc kBTlt0 is the translational (Stokes-Einstein) diffusion coefficient of an isolated Brownian sphere. In the data analysis gained from a typical EWDLS set-up such as the one sketched in Fig. 10.24, one conveniently decomposes the scattering vector into its components parallel and perpendicular to the wall, qk and q?, respectively. The first cumulant for the translational motion of a Brownian sphere near the glass wall is then given by C \u00bc q2k D0 k D E j \u00fe q2? \u00fe j2 4 D0 ? j; \u00f010:116\u00de where D0 k\u00f0z\u00de \u00bc kBTltk\u00f0z\u00de and D0 ?\u00f0z\u00de \u00bc kBTlt?\u00f0z\u00de, and h ij denotes a j-dependent weighted average of the z-dependent diffusion coefficients over all sphere - glass wall separations z, for a given evanescent wave penetration parameter j"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure11-1.png",
        "caption": "Fig. 11. 3D line tracking task using a car-like mobile manipulator (Proposed scheme).",
        "texts": [
            " The initial conditions are chosen as 1 2 3 [ (0) (0) (0) (0) (0) (0)] and (0) 0[0 0 0 0 / 2 / 2] T p p T x y \u03c6 \u03b8 \u03b8 \u03b8 \u03b3\u03c0 \u03c0= =\u2212 for which the end-effector position is initially located at ( , , ) (2,0,1).X Y Z = The steering wheel is allowed to change within the range of / 3 / 3.\u03c0 \u03b3 \u03c0\u2212 \u2264 \u2264 Virtual Joint Method for Kinematic Modeling of Wheeled Mobile Manipulators 1067 (a) Tracking error norm. (b) Manipulability. (c) \u03c6 and .\u03b3 (d) 1v q and 2 . v q (e) Joint angles of arm. (f) Velocity inputs of arm. Fig. 12. Results of a 3D line tracking task (Proposed scheme): Case 2. As shown in Fig. 11, the given two line trajectories are well followed by the WMM. In both the cases, the mobile platform moves backward initially and then turns its heading direction as to ultimately face toward the final point. This is because the tracking trajectories are behind the initial posture of the WMM. As the WMM approaches the final point, the arm becomes stretched out. Fig. 12 shows the error and state of the WMM during the tracking of Case 2 trajectory. Although rather sharp rises of error are observed when the steering angle changes much, the tracking performance is in general satisfactory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure5-1.png",
        "caption": "Fig. 5 The mesh map of the gear pump gear teeth in cycle two",
        "texts": [
            " The element effect matrix of H~P e4 and H~P e10 is obtained through interpolating in the solving domain, and the new response matrix H~1 can be obtained through combining H~P e4 with H~P e10 , and the damage coefficient matrix can be expressed as follows: D~ \u00bc \u00bdd11; d12; d13; d14; d15; d16 T \u00bc \u00bd 0:018; 0:3126; 0:0195; 0:0538; 0:0672; 0:0336 T The component which is negative value can be set as 0, and the new damage coefficient matrix can be expressed as follows: D~1 \u00bc \u00bdd12; d15 T \u00bc \u00bd0:075; 0:023 T The precision error is set as n = 10, and the same pro- cedure is carried out again. Cycle 2: the wavelet finite element model of the gear pump gear is shown in Fig. 5. D~2 \u00bc \u00bdd22; d27 T \u00bc \u00bd0:1632; 0:5467 T Cycle 3: the meshing situation is shown in Fig. 6. Intact xi 9.793 23.236 57.294 86.664 112.753 176.475 Cracked xi \u2019 9.361 22.873 56.541 85.357 111.532 175.352 D~3 \u00bc \u00bdd32; d35 T \u00bc \u00bd0:5832; 0:8836 T The zones P e32 and P e35 are identified as cracked element, and the angle is less than the error, and the iteration ends, the precious location, and the identification of the depth is carried out according to the contour line method. The damage parameters are defined as D~3 \u00bc \u00bd0; 0:6326; 0; 0; 0; 0 T for zone P e32, which are put into expression (32), and the changing rate of the natural frequencies of the multi-cracked gear pump gear Dxr xr corre- sponding to D~3 is calculated, and the results are listed as follows: Dx1 x1 \u00bc 0:153 %; Dx2 x2 \u00bc 0:457 %; Dx3 x3 \u00bc 0:338 % The crack identification of the gear pump gear can be carried out according to the variety of the natural frequencies of the gear with multi-cracks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001572_s11071-013-1014-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001572_s11071-013-1014-5-Figure1-1.png",
        "caption": "Fig. 1 MQL deep hole drilling principle [3]: (a) configuration of MQL deep hole drilling machine; (b) BTA deep hole drilling principle",
        "texts": [
            " Further, the dynamic behaviors and instability rotational speeds of drilling shaft are investigated numerically. The final section is concerned with the concluding remarks. 2.1 Model of MQL cutting fluid in deep hole drilling Deep hole drilling using MQL is a novel machining method. In the drilling process, the cutting edge of drilling tools is optimally cooled and lubricated, and simultaneously, the chip is removed by using the highpressure and low-temperature air with extremely small quantities of oil. The working principle is illustrated in Fig. 1. In the previous researches, the hydrodynamic forces of cutting fluid in deep-hole drilling process are normally formulated as analytical forms based on Reynolds equation, and this assumption is validated by comparing with practical data used in the factory or experiment [6\u20139]. Based on a similar idea, to solve the hydrodynamic pressure function of cutting fluid in MQL deep hole drilling, the authors consider the compressible fluid Reynolds equation in oil/air feature arising from cutting fluid to satisfy the finite length drilling shaft 1 r2 \u2202 \u2202\u03b8 ( p\u03b43 \u03bc \u2202p \u2202\u03b8 ) + \u2202 \u2202z ( p\u03b43 \u03bc \u2202p \u2202z ) = 6\u03c9 \u2202(p\u03b4) \u2202\u03b8 + 12 \u2202(p\u03b4) \u2202t (1) where r is the drilling shaft radius; \u03b8 is the circumferential coordinate of drilling shaft; \u03bc is the dynamic viscosity of cutting fluid; \u03b4 is the thickness of air/oil mist, \u03b4 = c(1 + \u03b5 cos\u03d5); \u03b5 is the transverse displacement ratio of drilling shaft; c is radial clearance between the drilling shaft and deep hole surface; \u03d5 is the angle of drilling shaft whirling; p is the hydrodynamic pressure function of cutting fluid; t is time; \u03c9 is the rotational speed of drilling shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure1-1.png",
        "caption": "Fig. 1. Configuration of a differential-drive wheeled mobile manipulator.",
        "texts": [
            " Section 2 reviews the quasi-velocity based kinematics of mobile manipulators as described in [2]; Sections 3 and 4 address in detail our virtual joint method for kinematic modeling and the related inverse kinematics; and Section 5 illustrates some simulation results, showing how to apply the proposed method. This section reviews the quasi-velocity based kinematic modeling of WMMs addressed in [2]. The formulation introduced here has been the fundamental idea for the kinematic formulations of the reduction based method. Consider the mobile manipulator shown in Fig. 1, in which an n degrees-of-freedom (DOFs) robot arm is mounted on a differential-drive mobile platform. The end-effector pose of the WMM with respect to the fixed frame, \u03a30, is 0 0 0 ( ) ( ) (3) 0 1 e e e SE \u23a1 \u23a4 = \u2208 ,\u23a2 \u23a5 \u23a3 \u23a6 R q p q A (1) where 0 3 3( ) e \u00d7 \u2208R q denotes the rotation matrix of the end-effector frame, \u03a3 n , with respect to \u03a30, 0 ( ) e p q is the end-effector position, and 1 2 [ ] T p p nx y \u03c6 \u03b8 \u03b8 \u03b8=q 3n+ \u2208 is the configuration variable, consisting of [ ] T b p px y \u03c6=q related to the platform position, ( ), p p x y, and orientation, ,\u03c6 at p O and 1 2 [ m \u03b8 \u03b8=q ] T n \u03b8 related to the manipulator configuration",
            ", the one for mobile platform and the other for its on-board arm), the proposed approach describes the motion of a WMM only by an equivalent single articulated manipulator. We shall consider two representative types of mobile platforms, from a simple differential-drive type to a more complicated car-like type in order to apply the proposed modeling approach - however, in principle, the applicability of the approach is not limited to these. 3.1. Differential-drive mobile platforms WMMs with differential-drive mobile platforms as shown in Fig. 1 can be physically transformed into ordinary articulated manipulator systems as shown in Fig. 2, where an imaginary revolute joint is anchored at the center of the wheel axle, Op, and an imaginary prismatic joint is attached atop the revolute joint. The structure of the transformed system exhibits the velocity kinematics identical to that of the original mobile manipulator. A systematic way to carry out the transformation is summarized as follows: Step 1: Fix a virtual revolute joint to ground at the current center position of the axle of the driving wheels, Op"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000882_0954405414564405-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000882_0954405414564405-Figure2-1.png",
        "caption": "Figure 2. Illustration of the gear hobbing machine axes.",
        "texts": [
            " Two spindles can be used as the tool axis and the workpiece axis. To synchronize these two axes, a Master-Slave EGB is used. In a Master-Slave EGB, a synchronous pulse is produced from the feedback pulse of the position detector attached to the tool axis (master axis) in the motor control, and the workpiece axis (slave axis) rotates with the pulse. The feedback pulse from the master side to the slave side is transmitted between spindle amplifiers. An illustration of the gear hobbing machine axes is shown in Figure 2, in which B is the hob spindle, C is the workpiece axis, X is the radial feed axis, Y is the tangential feed axis, Z is the axial feed axis and A is the hob installation angle adjustment axis. A theoretical control model of the Master-Slave EGB is illustrated in Figure 3, where RB is the control signal input to the spindle motor driver, KaB(A=V) is the current amplifier gain of the spindle motor driver, KtC(Nm=A) is the spindle motor\u2019s torque constant, TfB(Nm) is the nonlinear disturbance of the Hob (which is caused by any cutting force and nonlinear friction), JB(kgm 2) is the equivalent inertia, BB(kgm 2=s) is the viscous damping, KB is the transmission ratio of the reduction gearbox and uB is the actual position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002770_0142331219897186-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002770_0142331219897186-Figure2-1.png",
        "caption": "Figure 2. Engagement geometry.",
        "texts": [
            " The motion differential equations of the missile and the target are, respectively, given by _xM =VM cosa, _yM =VM sina, _a = aM=VM , _aM =(u aM=)tM , 8>< >: \u00f035\u00de where (xM , yM ) is the position coordinates of the missile in inertial ordinate system, aM presents the missile lateral acceleration, and tM denotes the missile autopilots time constant. _xT =VT cosb, _yT =VT sinb, _b = aT=VT , _aT =(v aT=)tT , 8>< >: \u00f036\u00de where (xT , yT ) is the position coordinates of the target in inertial ordinate system, aT presents the target lateral acceleration, and tT denotes the target autopilots time constant. This section consider the geometry of planar interception as provided in Figure 2, which is borrowed from (Sun and Liu, 2018c), where (X ,Y ) indicates position coordinates. Accordingly, the relative kinematics equation of the missiletarget intercept system is obtained as follows, and some related notations will be used to simplify the upcoming description (see Table 2) _r=VT cos (b u) VM cos (a u), r _u=VT sin (b u) VM sin (a u), _a= u=VM , _b= v=VT , 8>< >>: \u00f037\u00de Taking the time derivative of the LOS angle rate _u, and bearing (37) in mind, we get \u20acu= 2_r r _u cos (a u) r u+ cos (b u) r v: \u00f038\u00de To accomplish the goal of intercepting maneuverable aircraft, the relative distance r will be converged to zero by adjusting the missile control vector u",
            " The sampling period of the timetriggering scheme is given as h= 0:02s: Finally, a small pumping signal n(t)= sin5 (t) cos (t)+ sin5 (2t) cos (0:2t) is added to the control input in the first 10 seconds to ensure the persistent excitation effect. The PETADP algorithm presented in this paper is used to the missile-target interception system, then the following Figures 3\u201311 is obtained. As show in Figures 3\u20134, the LOS angle rate _u is converge to the neighbourhood of zero, and the range rate along the LOS is always negative, that is, _r\\0. Obviously, the maneuvering target is successfully captured by the missile in Figure 2 since both _u and _r satisfy the capture criteria in Remark 5. Figure 5 and Figure 6 reflect the lateral accelerations aT and aM . Figure 7(a) and (b) show the relative distance r between the missile and the target under the timetriggered scheme and the PET scheme, respectively. It is clear that relative distance r\\3 in Figure 7(b) while r\\10 in Figure 7(a), which illustrated that the the periodic eventtriggered ADP technique can implement more precise control for missiles. Figure 8 reflects the trajectory of engagement between the missile and the target, and the trajectories of W\u0302c is given in Figure 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002290_b15961-60-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002290_b15961-60-Figure1-1.png",
        "caption": "Figure 1. Exemplary illustration of the cycle time in press hardening.",
        "texts": [
            " Research aims are: the reduction of process cycle time, further enhancement of mechanical properties of hot formed metal sheets, further reduction of wall thickness and a general reduction of amount of energy used per component in its manufacturing. After examining mass production in the automotive stamping plant including existing problems, the project partners have jointly developed a representative demonstrator.To enable an easy transfer of project results into production, the demonstrator was designed to represent typical properties of series-production parts. The cycle time in hot sheet metal forming (press hardening) is determined to as much as 30% (see Figure 1) by the cooling time (holding time of closed die after forming before re-opening for part extraction). So it was assumed that, through an optimized cooling system manufactured by laser beam melting, cycle time of the hot forming process can be reduced significantly. Furthermore, it was assumed that by increasing the cooling rate, an improvement of the component\u2019s strength can be achieved. Subsequently, through the improved mechanical properties, a further reduction of wall thickness is getting possible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure7-1.png",
        "caption": "Fig. 7: Plate-to-plate welding case",
        "texts": [
            " 1 , , , ,( ) ( ) ( )s m m m tcp s tcp s m tcp s tcpU t U U t U t (5) where , ( )m m tcpU t , s mU can be obtained as in the tightly couple cooperation situation; , , ( )m tcp s tcpU t is time-varying, so (3) cannot be used here. It is often calculated as the relative motion defined by the task requirements of real-life application. More details can be found in the loosely coupled cases as follows. The master-slave robot kinematic cooperation approach can be widely used in welding applications. In this paper, the cooperated welding models are derived separately and applied into three typical welding scenarios namely the plate-to-plate, the tube-to-plate, and the tube-to-tube cooperated dual-robot welding systems. In Fig. 7 to 9 the red dashed line represents master robot kinematic chain and yellow dashed line represents salve robot kinematic chain. These two chains constitute a closed kinematic chain. The plate-to-plate welding is a typical kind of tightly coupled cooperation multi-robot system, which requires the relative pose of end tools between dual robots remaining unchanged in the welding process, as Fig. 7 shows. Note that the signal t in the figure means tcp for simplification. The tube-to-plate and tube-to-tube welding cases are two typical kinds of loosely coupled robot coordination, which requires a gripping robot and a welding robot, as Fig. 8 and 9 shows. The gripping robot is treated as the maser and used to catch and transfer the workpiece into the expected pose with the ship welding constraint, meanwhile the welding robot do the welding job. According to (5), m sU is obtained by robot calibration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003786_rteict49044.2020.9315633-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003786_rteict49044.2020.9315633-Figure1-1.png",
        "caption": "Fig 1: Piezoelectric Effect.",
        "texts": [
            " From the Greek piezo, meaning press or crush, and electron, meaning amber, the antique base of the 20 20 I nt er na tio na l C on fe re nc e on R ec en t T re nd s on E le ct ro ni cs , I nf or m at io n, C om m un ic at io n & T ec hn ol og y (R T E IC T ) | 9 78 -1 -7 28 1- 97 72 -2 /2 0/ $3 1. 00 \u00a9 Authorized licensed use limited to: Western Sydney University. Downloaded on June 15,2021 at 08:43:24 UTC from IEEE Xplore. Restrictions apply. 379 electric charger. French scientists Jacques and Pierre Curie discovered piezoelectricity in 1880. The piezoelectric effect as shown in Fig. 1 without pressure the output voltage is zero and as pressure applied current and voltage is produced due to the direct electromechanical interaction between mechanical and electrical conditions in crystalline objects without transfer equilibrium. Anirudh et.al [2] described work is about identifying, optimizing the piezoelectric effect & designing the ceramic tile which has the optimum piezoelectric effect. The pressure caused by foot stepping into the ceramic tile is one of the sources of vibration that is used as a trigger to produce energy using piezoelectric materials which is contained in the tile the outcome of the prototype of ceramic design is characterized including shrinking measurement density fracture strength test and resistivity material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002850_s00170-020-05095-2-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002850_s00170-020-05095-2-Figure4-1.png",
        "caption": "Fig. 4 Temperature field",
        "texts": [
            " h2 \u00bc 19:07 0:6 0:18 \u00bc 63:57 \u00f025\u00de Because the left bearing sleeve and spindle housings are fixed, the heat dissipation mainly depends on the air flow. Assuming that the air velocity is 0.35 m/s, for the left bearing sleeve, the laminar heat transfer coefficient can be calculated as follows. h3 \u00bc 41:58 0:0259 0:22 \u00bc 4:89 \u00f026\u00de All the boundary conditions of convective heat transfer are applied to the finite element model of the spindle. (4) Simulation Assuming that the ambient temperature is 20 \u00b0C and the simulation time is 1 h, the temperature field of the spindle can be simulated using the software of ANSYS. Figure 4 shows the temperature field of the spindle. It can be seen from Fig. 4 that the maximum temperature rise occurs in the middle angular contact ball bearings, and the temperature rise is 37.805 \u00b0C. However, the temperature rises of the right diaphragm coupling, and the grinding wheel shaft is only 0.124 \u00b0C due to the optimization of the thermal contact resistance. Figure 4 also shows that the heat transferred to the grinding wheel shaft from the left diaphragm coupling is also very small, that is to say, the optimization method of heat transfer can effectively reduce the temperature rise of the key component; it is useful to control the routes of heat transfer in high-speed spindle. Using the temperature field, the thermal deformation of the spindle can be analyzed by thermal-structure coupling analysis. Because the middle bearings are fixed, however, the other two sets of bearings can move freely along the axial direction, and the constraints of these three sets of angular contact ball bearings are applied to the finite element model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001025_2015-01-1477-Figure36-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001025_2015-01-1477-Figure36-1.png",
        "caption": "Figure 36. Scratch mark location and orientation from roll 2-5/8 ground contact (top) and plan view depicting scratch marks with velocity vector during ground contact (bottom).",
        "texts": [
            " As the component of perimeter velocity due to roll approaches the translational speed of the vehicle, the relative contribution of the direction of the CG velocity is decreased and the contribution resulting from roll velocity is increased. Additionally, the small effects of pitch and yaw rate on perimeter velocity will become more significant when the roll rate contribution approaches the CG velocity. For the left fender ground contacts, the perimeter speeds due to roll velocity were determined to be within 4 mph of the CG velocity during the 3rd and 4th revolutions. This condition resulted in abrasion patterns that were not parallel to the orientation of the CG velocity vector. Figure 36 and Figure 37 illustrate the scratch marks resulting from the roll 2-5/8 and roll 3-5/8 ground impacts located on the left front fender, hood, and left A-pillar. Scratch marks are first shown projected onto a CAD model of the pickup in the top portion of the figures. The lower illustration depicts the left front fender and hood with portions of the CAD model cut away. The model was oriented in the position at the ground contact, and the vehicle velocity vector is shown by the orange arrow. The scratch marks can be observed to be at a greater angle than the velocity vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003889_gt2014-25223-Figure20-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003889_gt2014-25223-Figure20-1.png",
        "caption": "FIGURE 20: 2D DFT diagram : eighth burst",
        "texts": [
            " These frequencies are analogous to sidebands as if vibration at \u03c94 r was modulated by rotating frequency [17]. It can be shown that non-linearity can explain the occurrence of sidebands. A detailed description of sidebands is beyond the scope of this paper and will be discussed in future publications. During the bursts, speed decreases, thus leading to a sweeping effect of mode frequencies by means of engine order excitation. At the final burst, engine order EO79 coincides with the first family mode with 3ND, as shown in Figure 19b. The 2D DFT diagram shown in Figure 20 highlights the nodal diameter content of the impeller response during the eighth burst. The third nodal diameter forward mode of the predominant first family is observed, along with several modes of higher families. At the same time, the first family modes on the casing with a 1ND forward traveling wave, as well as 2ND and 3ND backward traveling wave, produce the highest amplitudes. The terms forward and backward are correlated with the rotating direction (clockwise). Let\u2019s note that on the casing, the modal participation of higher families is imperceptible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001524_asjc.1157-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001524_asjc.1157-Figure8-1.png",
        "caption": "Fig. 8. Directed networks.",
        "texts": [
            " First, we formulate the pinned nodes selection problem as an optimization problem in directed case. Since the network is directed, then L is un-symmetric. Define L\u0304 = 1 2 (L + LT ) as symmetric part of L. Similar to the method in P3, correspondingly, a sub-optimal problem formulation assumes the form \u00a9 2015 Chinese Automatic Control Society and John Wiley & Sons Australia, Ltd (P4) \u2236 minimize s (10) subject to [ sI W\u0304 W\u0304 sI ] \u2ab0 0, \ud835\udf061(L) > 0 W\u0304 = I \u2212 \ud835\udf16L \u2212 k \u22c5 \ud835\udf16diag(\u0393), \ud835\udefe(i) \u2208 {0, 1}, \u0393\ud835\udfcf = p. (11) Example 5. Consider a directed network as Fig. 8a shows. Note that the network is strongly connected, thus each node can reach all the remaining nodes in the network. Fig. 9 demonstrates that the whole network can reach consensus if the root node is controlled. Different from the results in undirected case, node 1 who has the least PageRank and BC value, the second least CC value, and not the largest in-degree and out-degree is still chosen as the suboptimal pinned node for p = 1 by solving P4. As Fig. 9 shows, there are not any single indexes approaching the optimal strategy obviously. It is a challenge to find any index to characterize the node optimal properties. We further investigate the problem in a weakly connected directed network. Example 6. In Fig. 8b, only node 1, 2 and 3 can reach all the other nodes in the network, which implies that they are root nodes. As Fig. 10 shows that controlling node 6 can not drive all the nodes\u2019 state reach on desired consensus. Obviously, the other strategies except the optimal one lead to poor performance in Fig. 11. Comparing the convergence time of the weakly connected network with that of the strongly connected network under optimal strategy, we further find that the former converges faster than the latter, which is contrary to the fact that the weaker the network connectivity, the faster the convergence rate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.35-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.35-1.png",
        "caption": "Figure 2.35 (A) Carbon filled rubber electrode for TENS. Rubber pads of different formats, lead wire connections and sizes with adhesive patch for fixing. (C) Schematic diagram of the nerve fiber stimulation with the electric stimulus applied through surface electrodes.",
        "texts": [
            " When the target is located distant from this current path, or is protected by a high impedance tissue, like the dura mater, a monopolar work electrode and a return electrode can be used to increase the spatial coverage of the stimulus. Some limitations of the parallel bipolar stimulating electrodes are the difficulty in placing the two tips in the desired orientation, for instance, on the tissue slice for optimal tissue activation and the relatively large size of the electrode tips. Concentric bipolar electrodes possess the lowest stimulus artifacts and are easier to orient in the tissue, since they have no directionality. Figure 2.35A shows an example of surface, noninvasive stimulating electrode: flexible stimulating electrode for TENS. Instead of metal, it is made of silicone rubber impregnated with carbon particles; it is flexible, which facilitates its adjustment to body shape. The coupling with the skin is made via specific electrolytic gel (hydrogel with C1) and elastic bands help positioning the electrode in place; alternatively, the bottom surface of the electrode is covered with hypoallergenic adhesive for fixing. Stimulating pads typically are reusable and adhesive, made to stick directly on the skin with the help of a patch (Figure 2.35B). TENS electrodes are available in multiple sizes and features, according to the use. It is important to assure a uniform electrical current distribution over the contact surface to avoid skin burning and other chemical reactions. Figure 2.35C shows a schematic diagram of the functioning of TENS. A pulse of electric current or voltage is applied between the surface electrodes, generating an electric field in the conductor volume formed by the tissues between the skin and the nerve fiber. Ionic current flowing between the two electrodes reaches the fiber membrane; if the current exceeds the threshold of excitability, the cell membrane depolarizes and the action potential is generated. TENS is used in therapeutic treatment of chronic and acute pain (arthritic or muscular) and motor rehabilitation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003635_s00170-020-06321-7-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003635_s00170-020-06321-7-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of honing principle and crosshatch pattern",
        "texts": [
            " The cylinder liner is the core component of energy conversion and power output of engines.When the engine operates, the piston reciprocates in the cylinder liner, and the cylinder liner is subjected to temperature changes and pressure [4]. Therefore, the machining accuracy requirement of the cylinder liner is very high. Due to the high accuracy and surface quality of the honing process, the last machining process of cylinder liner is the honing process commonly [5]. The principle of the honing process of the cylinder liner is shown in Fig. 1. During the honing process, the oilstone is attached on the honing head to rotate and reciprocate at the same time; thus, the spatial trajectory of the oilstone is spiral. The oilstone exerts pressure on the inner surface of the cylinder liner and the materials of the cylinder liner are removed by the interaction of abrasive particles and the inner wall of the cylinder liner. The crosshatch on the inner wall of the cylinder liner is honed by the movement of honing oilstone; the angle of the crosshatch is determined by the reciprocating and rotating speeds of the honing head"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000115_mrs.2019.8901102-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000115_mrs.2019.8901102-Figure1-1.png",
        "caption": "Fig. 1. Results of (3) for T = 50 (black) and 150 (yellow) seconds.",
        "texts": [
            " , N}, [K] , {0, 1, 2, \u00b7 \u00b7 \u00b7K}, min xi N\u2211 i=0 K\u2211 k=0 [\u00b5pik log\u00b5ik \u2212 \u00b5ik + 1]\u03b2\u2206k (3a) N\u2211 i=0 K\u2211 k=0 [p\u00b5pik log\u00b5ik \u2212 \u00b5pik + 1]\u03b2\u2206k = \u2212 log\u03b1 (3b) \u00b5ik = 1 + \u03c7a \u03b2[2\u03c7+\u2016xik\u2212xtk\u20162] \u2200i \u2208 [N ], \u2200k \u2208 [K] (3c) \u2016xik \u2212 xjk\u2016 \u2265 d1 \u2200i, j \u2208 [N ], \u2200k \u2208 [K] (3d) \u2016xik \u2212 xtk\u2016 \u2265 d2 \u2200i \u2208 [N ], \u2200k \u2208 [K] (3e) The value of parameter p is obtained through (3b). If (3b) is not satisfied, then the detection problem is not well-posed; there is not enough data to make any meaningful decision. Statements (3c), (3d) and (3e) serve to formally define \u00b5, and place inter-MAV collision avoidance, and MAV-target safety distance constraints, respectively. Solution to (3) results in a minimum for the upper bound on PMD, should the detectors spend time T arranged in this configuration. Figure 1 shows the results of a 2D numerical simulation in which (3) is solved for 3 detectors, constrained to stay at a radius of radius 3 m around their target. The values selected for the parameters, \u03c7 = 2.77 \u00d7 10\u22126 m2 and \u03b2 = 0.00583 counts per second (CPS), are experimentally determined for the Domino Neutron detector,3 while a = 1.7\u00d7104 approximates a 5\u00b5Ci source4 available. The upper bound on PFA is set at \u03b1 = 10\u22123, and distances d1 and d2 are set at 1.5 m and 0.5 m, respectively. The standard solver NLOPT had been used to solve the nonlinear optimization.5 Decision deadline T was picked at 50 and 100 seconds in two different solutions that are compared against each other in Figure 1. Interestingly, the optimal solution does not distribute all the detectors symmetrically around the target; rather, it places one very close to the target while keeping the other two trailing while satisfying all constraints. That behavior is observed consistently for different decision deadlines and source intensities. The nonlinear dependence of SNR to the distance between sensor and source dictates that having at least one sensor as close as possible is preferable to uniformly minimizing all sensor-source distances"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure6-1.png",
        "caption": "Fig. 6 Intersection of the driving pin with the worm axial plane",
        "texts": [
            " Based on this, the angular velocity of the spindle is: wg \u00bc 2 p ng \u00f015\u00de The peripheral speed of the spindle alongside the radial rm pitch circle is constant, that is: vm \u00bc wg rm \u00f016\u00de Due to the shifting of the worm shaft by half cone angle, along the ellipse path curve which is perpendicular to worm axis, the angular velocity is changing but the peripheral speed of the spindle is constant: wP \u00bc vm rPe \u00bc changing \u00f017\u00de During the manufacturing process of conical worms, the changing angular velocity causes changing angular rotation which then causes pitch fluctuation. In order to eliminate it, it is important for the path curve of the driving pin to be a circle path instead of an ellipse path on perpendicular to axis section. That is why the angular velocity fluctuation and pitch changing do not occur, namely: wP \u00bc wg \u00bc const: \u00f018\u00de As a consequence of this: rPe \u00bc rm \u00f019\u00de Figure 6 shows if we intersect the worm with an axial plane which is at the half of the lathe fork, then the path curve of the driving pin is an ellipse and the section of the driving pin is an ellipse curve as well. Thus, the ellipse cross section is moving on an ellipse path. Based on the CDE rectangular triangle, the major axis of ellipse is (Fig. 6): dcsen \u00bc dcs cos d1 \u00f020\u00de The minor axis of ellipse is: dcsek \u00bc dcs \u00f021\u00de In order to eliminate the angular velocity fluctuation, the moving ellipse has to be placed onto a radial circle path rm (Fig. 7). The displacement distance on the major radius of the ellipse is (Fig. 7): rhn \u00bc rne rm \u00f022\u00de The calculated displacement distance for the arbitrary ellipse P point is (Fig. 7): rhP \u00bc rPe rm \u00f023\u00de The polar equation of the ellipse section of the driving pin is (Fig. 7): xcs \u00bc rcs sin 8 1 ycs \u00bc rcsen cos 8 1 \u00f024\u00de Using the given equations above, we have prepared our own computer program (Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002540_ssrr.2016.7784286-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002540_ssrr.2016.7784286-Figure2-1.png",
        "caption": "Fig. 2. Three phases of traversing stairs",
        "texts": [
            " This assumption means that the robot can traverse the stairs while maintaining an inclination of the robot that is equal to that of the stairs. In addition, the conditions of (p, \u03b8s) are limited in that falling backward (described later) does not occur in the case the robot stretches all subtracks completely straight. This assumption physically means that the robot at least has a way to traverse the stairs. While a tracked robot traverses the stairs, there are roughly three phases of the motion (shown in Fig. 2), and each phase has different characteristic problems. In this section, we describe the definitions of the three phases and the problems that tend to occur in each phase. First, we divide the stair climbing motion into three phases, as follows: (i) Pitch-up phase: The phase starts when the robot touches the first step of the stairs and lasts until the inclination of the robot corresponds to that of stairs (Fig. 2-(i)) (ii) Normal climbing phase: The phase between (i) and (iii). During this phase, the inclination angle of the robot is equal to that of the stairs (Fig. 2-(ii)). (iii) Pitch-down phase: The phase during which the inclination angle of the robot starts decreasing and leads to the robot successfully reaching the new level (Fig. 2-(iii)). While the robot traverses the stairs, it is necessary to prevent the following four failure modes: 1) Slipping mode: The friction between the tracks and the steps is not sufficient to move the robot forward. 2) Falling backward: The robot body rotates backward and downward around the lowermost contact point to the stairs (flips over backwards). 3) Falling sideward: The robot body rotates sideward around the axis of the contact line of the robot\u2019s side (tips sideways as shown in Fig. 3). 4) Excessive shock: The contact shock exceeds the acceptable level for the robot (or surrounding environment) when it climbs over the final step of the stairs(as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002947_j.engappai.2020.103629-Figure25-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002947_j.engappai.2020.103629-Figure25-1.png",
        "caption": "Fig. 25. Definition of the body and the NED frames.",
        "texts": [
            "2) \u23a1 \u23a2 \u23a2 \u23a3 ?\u0307? ?\u0307? ?\u0307? \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a3 1 sin\ud835\udf19 tan \ud835\udf03 cos\ud835\udf19 tan \ud835\udf03 0 cos\ud835\udf19 \u2212 sin\ud835\udf19 0 sin\ud835\udf19 sec \ud835\udf03 cos\ud835\udf19 sec \ud835\udf03 \u23a4 \u23a5 \u23a5 \u23a6 \u23a1 \u23a2 \u23a2 \u23a3 \ud835\udc5d \ud835\udc5e \ud835\udc5f \u23a4 \u23a5 \u23a5 \u23a6 (F.3) ?\u0307?\ud835\udc01 \ud835\udc0d\ud835\udc01 = \ud835\udc09\u22121 [ \u2212\ud835\udf4e\ud835\udc01 \ud835\udc0d\ud835\udc01 \u00d7 ( \ud835\udc09\ud835\udf4e\ud835\udc01 \ud835\udc0d\ud835\udc01 ) +\ud835\udc26\ud835\udc01] (F.4) where, \ud835\udc29\ud835\udc0d = [\ud835\udc5d\ud835\udc5b \ud835\udc5d\ud835\udc52 \ud835\udc5d\ud835\udc51 ]T is the position of the UAV in the NED frame and \ud835\udc2f\ud835\udc01 = [\ud835\udc62 \ud835\udc63 \ud835\udc64]T denotes the velocity of the UAV, expressed in the body frame. Also, \ud835\udf4e\ud835\udc01 NB = [\ud835\udc5d \ud835\udc5e \ud835\udc5f]T denotes the angular velocity of the body frame with respect to the NED frame, expressed in the body frame. The body and the NED frames are shown in Fig. 25. Moreover, [\ud835\udf19, \ud835\udf03, \ud835\udf13] are Euler angles, m represents the mass and \ud835\udc02\ud835\udc0d \ud835\udc01 represents the DCM of the NED frame with respect to the body frame, calculated as (Zipfel, 2007) \ud835\udc02\ud835\udc0d \ud835\udc01 = \u23a1 \u23a2 \u23a2 \u23a3 c\ud835\udf03c\ud835\udf13 s\ud835\udf19s\ud835\udf03c\ud835\udf13 \u2212 c\ud835\udf19s\ud835\udf13 c\ud835\udf19s\ud835\udf03c\ud835\udf13 + s\ud835\udf19s\ud835\udf13 c\ud835\udf03s\ud835\udf13 s\ud835\udf19s\ud835\udf03s\ud835\udf13 + c\ud835\udf19c\ud835\udf13 c\ud835\udf19s\ud835\udf03s\ud835\udf13 \u2212 s\ud835\udf19c\ud835\udf13 \u2212s\ud835\udf03 s\ud835\udf19c\ud835\udf03 c\ud835\udf19c\ud835\udf03 \u23a4 \u23a5 \u23a5 \u23a6 (F.5) where s and c stand for sin and cos functions, respectively. Also, J is the inertia matrix, given by Zipfel (2007) \ud835\udc09 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 \u222b (\ud835\udc662 + \ud835\udc672)\ud835\udc51m 0 \u2212 \u222b \ud835\udc65 \ud835\udc67 \ud835\udc51m 0 \u222b (\ud835\udc652 + \ud835\udc672)\ud835\udc51m 0 \u2212 \u222b \ud835\udc65 \ud835\udc67 \ud835\udc51m 0 \u222b (\ud835\udc652 + \ud835\udc672)\ud835\udc51m \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 \u225c \u23a1 \u23a2 \u23a2 \u23a3 Jxx 0 \u2212Jxz 0 Jyy 0 \u2212Jxz 0 Jzz \u23a4 \u23a5 \u23a5 \u23a6 (F"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002423_jae-160067-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002423_jae-160067-Figure3-1.png",
        "caption": "Fig. 3. Configuration relationships between PVDF actuator and cantilever beam.",
        "texts": [
            " The magnitude of the light induced actuation strain of PVDF actuator can be expressed as: \u2212 S(t) = df3i V (t) Df e = df3i Vs ( 1\u2212 e \u2212 t \u03c41 ) + ( AP Cp \u2212 \u03b2 \u03bbDe d3iYa ) \u0394Ts ( 1\u2212 e \u2212 t \u03c4\u03b8 ) Df e (3) \u2212 Sd(t) = df3i Vd(t) Df e = df3i V (t0)\u2212 ( AP Cp \u2212 \u03b22\u03bbDe d3iYa ) \u0394Ts\u2212d + ( AP Cp \u2212 \u03b22\u03bbDe d3iYa ) \u0394Ts\u2212de \u2212 t \u03c4d Df e (4) where df3i is piezoelectric-strain constant of PVDF actuator, and Df e indicates the distance between two electrodes of PVDF actuator. The configuration relationships between PVDF actuator and cantilever beam are presented in Fig. 3. Based on the mathematical model of PLZT with coupled multi-physics fields and the constitutive model of cantilever beam, the deflection of this beam during illumination and light off phases can be written as follows, respectively. w(t) = \u2212 S (h+ ha) 2 habaYa La Y Ib ( L\u2212 La 2 ) = df3i [ Vs(1\u2212 e \u2212 t \u03c41 ) + (AP Cp \u2212 \u03b2 \u03bbDe d3iYa )\u0394Ts ( 1\u2212 e \u2212 t \u03c4\u03b8 )] Df e (h+ ha) 2 habaYa La Y Ib ( L\u2212 La 2 ) (5) wd(t) = \u2212 Sd (h+ ha) 2 habaYa La Y Ib ( L\u2212 La 2 ) = df3i [ V (t0)\u2212 (AP Cp \u2212 \u03b22\u03bbDe d3iYa )\u0394Ts\u2212d + (AP Cp \u2212 \u03b22\u03bbDe d3iYa )\u0394Ts\u2212de \u2212 t \u03c4d ] Df e (h+ ha) 2 habaYa La Y Ib ( L\u2212 La 2 ) (6) In order to facilitate the parameters calibration of the deflection equations of cantilever beam, we need to simplify Eqs (5) and (6)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001096_j.eaef.2014.02.002-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001096_j.eaef.2014.02.002-Figure4-1.png",
        "caption": "Fig. 4. Calculation of position and heading angle.",
        "texts": [
            "185 m, when the combine harvester turns, its speed (speed of the center of the crawlers) can be calculated by Eq. (3): vt \u00bc vs R R\u00fe 0:5925 ; (3) where vt is the vehicle\u2019s (center of the crawlers) speed (m/s) during turning, vs is the vehicle\u2019s speed (m/s) in a straight movement (outer crawler\u2019s speed) and R is the turning radius (m) calculated by Eq. (2). In the simulation program, the combine harvester\u2019s position and heading angle were calculated by the following method. When the combine harvester is turning, it moves in an arc, and the method shown in Fig. 4 was used. Assuming that combine harvester\u2019s current coordinates are (x0, y0) with a heading angle of 40 and yaw angular velocity of u; after a short period of time Dt, the combine harvester\u2019s coordinates (xt, yt) and heading angle 4t can be calculated by Eq. (4) and Eq. (5). yt xt \u00bc y0 x0 \u00fe dR jdj cos f0 sin f0 sin f0 cos f0 sinDf 1 cosDf (4) Df \u00bc ft f0 \u00bc u Dt; (5) where, according to Eq. (1), Eq. (2) and Eq. (3), yaw angular velocity u can be calculated by Eq. (6): u \u00bc vt R \u00bc \u00f00:0208l 0:0134\u00de 450:5\u00f0d\u00de 1:318 \u00fe 0:5925 : (6) In addition, when the robot combine harvester travels in a straight line, its position can be calculated by Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002239_1350650115592919-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002239_1350650115592919-Figure2-1.png",
        "caption": "Figure 2. Technical drawing of the Bochum test rig for large turbine bearings.",
        "texts": [
            " The key parameters of the bearing are listed in Table 1. The Bochum test rig has been designed to examine large turbine journal bearings in original size under practical operation conditions. It was first presented by Hopf1 and later also described by Hagemann, Kukla, and Schwarze.3,4 With a DC drive with 1.2MW the \u00d8 500mm shaft can be run at a top speed of 4000 rpm which results in a circumferential speed of 104.7m s 1. The lubricant used is turbine oil ISO VG 32. A technical drawing of the rig is shown in Figure 2. Following Glienicke\u2019s design,10 the shaft (2) is mounted in two support bearings (3). The test bearing (1) is centrally arranged between these support bearings, and placed in a rigid frame (9) which is connected to a pneumatic bellow (6). A maximum force of 1 MN can be exerted by pulling the test bearing against the shaft when the bellow is pressurized. The bellow, the traverse it is located on, and the support bearings are supported by the test rig body (4). The shaft is designed as a hollow shaft and is equipped with two piezoelectric pressure sensors and two capacitive distance sensors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003148_s40032-020-00591-6-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003148_s40032-020-00591-6-Figure1-1.png",
        "caption": "Fig. 1 Solid model of the rotorbearing system",
        "texts": [
            " The remaining part of the manuscript is organized as follows: section two presents the mathematical modeling of rotor-bearing system under consideration along with details of bearing film forces and transient unbalance excitations. Section three gives results and discussions, while some conclusions of the work are highlighted in section four. The rotor assembly under consideration consists of a rotor with two disks (namely turbine wheel (TW), and compressor wheel (CW)) and bearings in its simplest form. Figure 1 shows the three-dimensional model of system assembly. A uniform cross-section of the flexible rotor is considered, and disks are assumed as rigid. The FE beam model of the rotor-bearing system with eight elements is shown in Fig. 2 by considering the shear and bending effects. With a convergence study, it was found that an eight element model approximates consistently the first few natural frequencies. The disks are treated as rigid lumped masses placed at the nodes one and nine, and the floating ring bearings are placed at the nodes three and seven"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003106_978-3-030-43089-4_23-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003106_978-3-030-43089-4_23-Figure12-1.png",
        "caption": "Fig. 12: Untangling a double-coin knot.",
        "texts": [
            " The IK problem is to move the center of the hand to a specific position, and the workspace grid has 8,000 points. A huge number of disconnections are caused by pointwise resolution. After CSP resolution (Fig. 8, right), the number of disconnections drops to a minimal number (0.062% of reachable edges). This favorable result is likely because the Jaco has three continuous rotation joints. The next four figures show optimized results for single arms of the Boston Dynamics ATLAS (Fig. 9), the Rethink Robotics Baxter (Fig. 10), the NASA Robonaut2 (Fig. 11), and a leg of the JPL Robosimian (Fig. 12). None of these robots has continuous rotation joints. The kinematic structure and limits of each robot is different, with the ATLAS, Baxter, and Robonaut2 having an 508 K. Hauser Fig. 9. Optimized results for the arm of the ATLAS humanoid. Two views of the discontinuity boundary are shown. Fig. 10. Optimized results for the arm of the Baxter robot. Two views of the discontinuity boundary are shown. anthropomorphic elbow and 3-axis shoulder, but with different axes and joint limits. The Robosimian\u2019s joints are arranged so that each pair of 2 joints have intersecting axes, and each joint can rotate 360\u25e6",
            " It also has discontinuities the opposite side of the torso near the rear, where the arm can reach points either around the front or around the back of the torso. Robonaut2 has essentially the opposite problem, with discontinuities along the underside of the arm. Joint limits in the shoulder mean that its hand must move up and around to pass from elbow-back to elbow- Continuous Pseudoin ersion of a Multi ariate Function 509 Fig. 11. Optimized results for the arm of the Robonaut2. Two views of the discontinuity boundary are shown. Fig. 12. Optimized results for one limb of the Robosimian quadruped. Two views of the discontinuity boundary are shown. front configurations. Other small discontinuities are scattered about the torso, which are likely free space narrow passages caused by self-collision. Perhaps a denser sampling of configurations or workspace would help reduce these artifacts. Robosimian is an interesting case. Although its joints have a larger range of motion than each of the four other robots, its kinematics are quite different"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001811_s11431-015-5883-3-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001811_s11431-015-5883-3-Figure3-1.png",
        "caption": "Figure 3 (Color online) Ordered film model for TFL [14].",
        "texts": [
            " As demonstrated in Figure 2 [14], it has been commonly recognized that there are five stages with the decreasing of the lubricant film thickness, which are hydrodynamic lubrication (HL), TFL, boundary lubrication (BL) and dry friction. The film formation in TFL deviating from the EHL has been contributed to the ordered molecules at the solid interface, which can be well pictured as a molecular distribution model with three zones (adsorbed boundary layer, ordered layer, fluid layer) from the solid wall to the center of the gap [14,25], as shown in Figure 3. In TFL regime, the core concept is the ordered molecular layer nearby the solid wall, which is resulted from the emerged physical effect of the solid surfaces on the lubricants. The interaction between the solid surface and the lubricant molecules can be highlighted when the lubrication state is discussed at nanometer scale, leading to the \u201cabnormal\u201d film formation properties in between the two rubbing surfaces different from that of EHL. The influence of solid surface energy on the film formation has been investigated and analyzed by both theoretical and experimental methods since 1994 [26\u201328], indicating that high-energy surface will strongly influence on the action of lubricant molecules"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003470_j.jnnfm.2020.104406-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003470_j.jnnfm.2020.104406-Figure1-1.png",
        "caption": "Fig. 1. Schematic drawings of the rectangular microchannel and the two cylinders.",
        "texts": [
            " Also depending n the proximity to the heart, a smaller or larger blood vessel can be the ifference between a Re \u2264 1 flow or a Re = 3000 flow, hence the variety f flow regimes in our circulatory system is immense [41,55,115]. lood capillaries are particularly interesting: by being so small (<1 mm n diameter) they enhance the viscoelasticity of blood as explained in he Introduction. Inspired by these, a straight, 10 mm long, rectangular icrochannel was designed. The cross section is 2\ud835\udc3fc = 270 \u03bcm wide nd \ud835\udc3b = 100 \u03bcm deep, setting the channel\u2019s characteristic length-scale t \ud835\udc3fc = 135 \u03bcm (see Fig. 1). The confinement ratio of the microchannel s \ud835\udefc = \ud835\udc3b\u22152\ud835\udc3f \u223c 0.37. c Since we are using non-particulate blood analogues, the size of he cylinders needs to be such that discards the possibility of discrete nteraction with blood particles. Only this way can the assumption of continuous medium be valid hence not having to consider a flow of uspensions. RBCs are the largest particles in human blood, measuring \u201310 \u03bcm in diameter [116]. With that in mind, the minimum length of he cylinders was defined as ten times the average RBC size: 75 \u03bcm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001131_j.amc.2014.01.138-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001131_j.amc.2014.01.138-Figure2-1.png",
        "caption": "Fig. 2. The ortho\u2013parallel manipulator. The cylinders represent the revolute joints. The global basis vectors coincide with the basis vectors of 1st body (red) in this configuration. (For interpretation of the references to colour in this figure caption, the reader is referred to the web version of this article.)",
        "texts": [
            " This choice leads to the decomposition Irev\u00f0a; b\u00de \u00bc I1 rev\u00f0a; b\u00de \\ I2 rev\u00f0a; b\u00de; Ij rev\u00f0a; b\u00de \u00bc hk j1 ; kj2 ; kj3 ; kj4 ; kj5 ; kj6 ; kj7i; \u00f08\u00de where the polynomials kij are given in Appendix A. We need also the decomposition of Irev\u00f0a; a\u00de for a \u00bc \u00f01;0;0;0\u00de and \u00f0g1;g2;v\u00de \u00bc \u00f0e2; e3; e1\u00de. which gives Irev\u00f0a; a\u00de \u00bc I1\u00f0a\u00de \\ I2\u00f0a\u00de \u00bc ha2 0 \u00fe a2 1 1; a2; a3i \\ ha0; a1; a2 2 \u00fe a2 3 1i: \u00f09\u00de Let us investigate an ortho\u2013parallel manipulator composed of three bars connected by revolute joints and an arm endeffector, see Fig. 2. This leads to the following system: w1 \u00bc \u00f0e2;R1e1\u00de; w2 \u00bc \u00f0e3;R1e1\u00de; w3 \u00bc \u00f0R 1e1;R2e3\u00de; w4 \u00bc \u00f0R 1e2;R2e3\u00de; w5 \u00bc \u00f0R2e1;R3e3\u00de; w6 \u00bc \u00f0R2e2;R3e3\u00de; w7 \u00bc h\u00f0a\u00de; w8 \u00bc h\u00f0b\u00de; w9 \u00bc h\u00f0c\u00de; Iw \u00bc hw1; . . . ;w9i: The position in Fig. 2 corresponds to the ideal I init \u00bc I a \u00fe Ib \u00fe I c \u00bc ha0 1; a1; a2; a3i \u00fe hb0 1= ffiffiffi 2 p ; b1; b2; b3 \u00fe 1= ffiffiffi 2 p i \u00fe hc0 1= ffiffiffi 2 p ; c1; c2; c3 \u00fe 1= ffiffiffi 2 p i: In order to choose the correct components in the decompositions (8) and (9) we now check that V\u00f0Ia\u00de V\u00f0I1\u00f0a\u00de\u00de V\u00f0Ia \u00fe Ib\u00de V\u00f0I1 rev\u00f0a; b\u00de\u00de V\u00f0I b \u00fe I c\u00de V\u00f0I1 rev\u00f0b; c\u00de\u00de: Hence the appropriate component of Iw is given by I \u00bc I1\u00f0a\u00de \u00fe I1 rev\u00f0a; b\u00de \u00fe I1 rev\u00f0b; c\u00de: \u00f010\u00de Let us also note that since a2 \u00bc a3 \u00bc 0 we can after substitutions interprete that I is an ideal in the ring A \u00bc Q\u00bda0; a1; b0; b1; b2; b3; c0; c1; c2; c3 : Hence we consider the constraint variety V \u00bc V\u00f0I\u00de as a subset of R10 and as usual we denote by f the map corresponding to the ideal I "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of gear",
        "texts": [
            " (3) Then the damage element corresponding to dj [ 0 is disposed, and the element will be refined, and the effect matrix H~ is computed again, through iteration the new damage matrix is obtained, this procedure is repeated and ended until the size of the refined element satisfies the predefined precision, and the cracks will be found out based on the damage coefficient finally. (4) The depths of the cracks are identified. The damage corresponding to coefficients the top three order natural frequencies will be computed and used as input parameters of the contour lines; and the contour lines corresponding to the every order modal frequency, and the intersections of the contour lines are the location and depth of the crack. The schematic diagram of gear with two cracks is shown in Fig. 2. The modulus of gear is 0.4, the number of teeth is 11, the pressure angle is 20 . The manufacturing material of the gear is C45, the elastic modulus of it is 200 GPa, the Poisson ratio of it is 0.3, and the density of it is 7,859 kg/ mm3. Based on the blind source separation with the particle swarm optimization algorithm, the natural frequencies are obtained. The natural frequencies of intact and cracked gear pump gear are listed in Table 1. The part of the gear pump gear tooth is divided into 11 zones, and the wavelet finite element division is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.6-1.png",
        "caption": "Figure 3.6 Permanent magnet brush motor cross-section",
        "texts": [
            " System Components 39 Beginning in 1930, a powerful permanent magnet material, Alnico (aluminum, nickel, cobalt) was created that has, to a large extent, replaced the wound stator poles and resulted in the development of today\u2019s permanent magnet brush motor. In the 1960s a relatively inexpensive permanent magnet material, Hard Ferrite, a compound of various ceramics and iron oxide was developed, helping to create a class of low cost brush motors. A cross-section of a typical four pole permanent magnet brush motor is shown in Figure 3.6. This is illustrated in Figure 3.7. Regardless of whether the field is created by current-carrying coils mounted on salient poles or by permanent magnets, armature control assumes that the field is constant and all control is dependent on control of the armature current. 40 Electromechanical Motion Systems: Design and Simulation The relation between the applied current and the resulting torque is expressed as: T = Kt Ia (3.3) where Kt is the torque constant in N m A\u22121, Ia is the applied current in A and T is the resulting torque in N m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure23-1.png",
        "caption": "Figure 23. Speed field contours for 800 rpm.",
        "texts": [],
        "surrounding_texts": [
            "Table 2 and the graph (Figure 19) show the calculations of the maximum speed of the velocity fields for a fixed drum diameter of 650 mm and different values of the drum rotation speeds. The diameter and rotation speed of the beater are equal to the initial values. [22] ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0000041_s11668-019-00763-2-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000041_s11668-019-00763-2-Figure1-1.png",
        "caption": "Fig. 1 The actual (a) and schematic (b) lower arms of the MF399 tractor",
        "texts": [],
        "surrounding_texts": [
            "In the present study, the dimensions of the lower arms of 399 and 285 Massey Ferguson tractors were measured. These were used to simulate a three-dimensional mechanism of lower arms in SolidWorks 2013, a commercial software. The real and schematic lower arms of MF399 and MF285 tractors are shown in Figs. 1 and 2, respectively. Then, the CAD simulation files in order to build the FEM were imported into ANSYS V15 Software. Mechanical properties of steel (St 37) which was used to build the lower arms are presented in Table 1. Static Analysis The finite element models were meshed using SOLID187 that is a three-dimensional solid element and has 8 nodes with three degrees of freedom for each node. The models consisted of 614 elements and 16,318 nodes for the lower arm of MF399 tractor and 530 elements and 99,089 nodes for the lower arm of MF285 tractor (Fig. 3). Meshing of models was performed by the free method. The boundary and displacement conditions were applied to arm junction points to the tractors. The values of loads applied to the lower arms of MF399 and MF285 tractors were obtained from field testing under different working conditions. The tensile force testing was performed in order to determine maximum value of forces on the lower arms and its direction while working with chisel, furrower equipment and drill planter. The lower arms supply draft forces which are applied in parallel with the forward direction of tractor. The tensile load of chisel plow and furrower equipment during tillage operation and drill planter while planting were measured at the maximum depth of working (Table 2) and were considered as input loads for analyzing of the model in ANSYS software. In static analysis, force was applied to the lower arms and after analyzing the model, the equivalent tension stress on the lower arms was calculated based on Von-Mises theory. To avoid of the parts failure, the maximum stress must not exceed of the material yielding strength. Under static loading, the factor of safety (FS) is obtained from yield strength (ryp) divided by maximum working or allowable stress (rall) applied on the lower links (Eq 1). In a proper design, the FS must be greater than one, indicating that the structure can work and remain healthy in uncertain loading conditions [21]. F:S: \u00bc ry rall \u00f0Eq 1\u00de Modal Analysis Modal analysis is necessary in the structural and industrial components designs that can resonate. Small loads at resonance frequency can result in induced deformation and damage in the structures. Therefore, mechanical parts must be designed as far as possible away from the resonance Table 1 Mechanical properties of steel (St 37) Density, q (kg/ m3) Modulus of elasticity, E (GPa) Poisson\u2019s ratio, t Yielding strength, ryp (MPa) Ultimate tensile strength, rut (MPa) 7860 200 0.3 198 235 Table 2 Equipment properties and the maximum loads applied to lower arms Implements Number of unit Maximum depth (cm) Tensile load (KN) Furrower 4 24 8.64 Chisel plow (sweep) 9 24 9.71 Drill planter 6 12 14.22 frequency range. Fluctuation in natural frequency of the structure increases the vibration amplitude and results in the failure and fracture of the component. Modal analysis is used to determine the value and mode shapes of the natural frequencies [22]. The natural frequency of the structure depends on the shape, material and supports of the structure. However, the amount and type of loads can affect the natural frequency. In this study, the SOLID 182 and 185 elements were used for modal analysis of the parts. To determine the characterization of materials including modulus of elasticity, Poisson\u2019s ratio and density of steel (St 37) were used (Table 1). The models were meshed, boundary conditions were applied in the lower arms connecting to the tractor, and desired nodes were binding in all directions. Then, the modal analysis of the lower arms of MF399 and MF285 tractors was performed with regard to the first 5 natural frequencies in range of 0\u20135000 Hz. Fatigue Analysis The effect of cyclic loads applied on structures creates cracks and finally results in part fracture, while the magnitude of reversing stress is lower than the yield stress of the structure. This phenomenon is called fatigue due to applying cyclic loads. In ANSYS software to perform a fatigue analysis under intermittent loads, firstly the exerted stresses in structures under cyclic loads must be determined. So, before any fatigue analysis, statistical analysis should be performed. Then according to stress contours, the critical nodes with maximum stresses must be detected and after fatigue analysis can be surveyed on the critical nodes. In order to obtain the endurance stress (re) of the steel material in the lower arms, fatigue analysis was performed on the model by applying 1.5 million loading cycles. For this purpose, the SOLID 185 element was used. It is a three-dimensional and 4-node element, and each node has three degrees of freedom. Characterization of materials, mesh and applying boundary conditions was conducted the same as those considered for static analysis. The Soderberg equation (Eq 2) was used to calculate factor of safety (FS) in fatigue analysis [23]. 1 F:S: \u00bc rave ry \u00fe K rr re \u00f0Eq 2\u00de where rave is average stress, rr is reversing stress, and K is geometric stress concentration factor. In the above equation, rave and rr were obtained by Eqs 3 and 4. In these equations, rmax was the maximum stress and rmin was considered as zero. rave \u00bc rmax \u00fe rmin 2 \u00f0Eq 3\u00de rr \u00bc rmax rmin 2 \u00f0Eq 4\u00de"
        ]
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure3-1.png",
        "caption": "Fig. 3. Kinematic structure of a car-like wheeled mobile",
        "texts": [
            " Another advantage of the virtual joint based modeling is that the dynamics of WMMs can be formulated more conveniently without consideration of the nonholonomic constraints. If the virtual joints are used, the resulting dynamic equation becomes simply expressed just like that of a conventional holonomic manipulator in a relatively easy manner. 3.2. Car-like mobile platforms WMMs with car-like mobile platforms, in contrast to the cases with differential-drive mobile platforms, have varying forms of nonholonomic constraints. For systems with front-drive and -steering wheels as shown in Fig. 3(a), three possible nonholonomic constraints are (C1) sin cos 0 (C2) sin cos 0 (C3) sin( ) cos( ) 0, p p m m m m x y x y D x y \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03b3 \u03c6 \u03b3 \u2212 = \u2212 + = + \u2212 + = (10) where (C1) implies the no-slip condition in the sideway direction of the rear wheels, (C2) implies the same constraint as (C1) but expressed in different variables, Hyunhwan Jeong, Hyungsik Kim, Joono Cheong, and Wheekuk Kim 1062 and (C3) implies the no-slip condition in the sideway direction of the front wheels. The directions of the quasivelocities, gi, if obtained based on (C1), can be exactly the same as (5), like the case with differential-drive mobile platforms. Therefore, the virtual joints for a carlike mobile platform can be placed following the same steps in the previous subsection. For instance, the mobile manipulator shown in Fig. 3(a) is transformed into the equivalent manipulator system of Fig. 3(b). The velocities of the virtual joints for such systems that realize the quasi-velocities are determined as 1 2 sin and cos f v v f v q q v D \u03c6 \u03b3 \u03b3= = = , (11) where 1v q and 2v q again denote the velocities of the revolute and prismatic joints, respectively; and vf and \u03b3 are the control variables, representing the driving velocity and the steering angle in the front wheel. The particular definition in (11) is chosen so as to meet the following two requirements: equivalence of the velocity kinematics, and preservation of the nonholonomic constraint. First, to show the equivalence of velocity kinematics, let us compare the velocities at Om between the original and the transformed systems. From Fig. 3(b), the translational velocity at Om is 1 2 sin cos sin sin cos cos cos( ), m v v f f f x Dq q v v v \u03c6 \u03c6 \u03b3 \u03c6 \u03b3 \u03c6 \u03c6 \u03b3 = \u2212 + = \u2212 + = + (12) 1 2 cos sin cos sin cos sin sin( ) m v v f f f y Dq q v v v \u03c6 \u03c6 \u03b3 \u03c6 \u03b3 \u03c6 \u03c6 \u03b3 = + = + = + , which are the same attained for the original system of Fig. 3(a) (Refer to [1] for verification). Since the rotational velocity of the platform is by definition 1 , v q \u03c6= the equivalence of the velocity kinematics between the original and the transformed systems has been verified. The preservation of nonholonomic constraints can be checked simply by substituting (11) and (12) into (10), which gives that (C2) and (C3) are satisfied, while (C1) is satisfied by construction. It is worth mentioning that mapping a connected domain of ( )fv\u03b3 , into 1 2 ( ) v v q q, via (11) produces disjoint images"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure2-1.png",
        "caption": "Fig. 2 Modeling of spiroid worm gear drives",
        "texts": [
            " Our objective was to determine the geometrical shaping of the driving pin in order not to have pitch fluctuation [2]. In technical practice, conical worm surfaces, which can be used in many different ways, are most widely applied as an active surface of conical worms [3]. The conical worm and crown wheel paired spiroid drives can be used for example as jointless drives of robots and tool machines. We can get jointless drives by simply shifting (setting) the worm into an axial direction [4, 5]. The cog surface of conical worms of the spiroid drives (Fig. 2) can be attributed in the same way as that of the cylindrical worm, but besides the axial shifting of the tool, a radial feeding must be done as well, depending on the conicity of the worm [6, 7]. Different\u2014evolvent, Archimedean and convolute\u2014helical surfaces [10, 11] can be defined in case of spiroid worm surfaces similar to the plain cylindrical worms. I. Dud\u00e1s : S. Bodz\u00e1s : Z. M\u00e1ndy Department of Production Engineering, University of Miskolc, 3515 Miskolc, Egyetemv\u00e1ros, Hungary I. Dud\u00e1s e-mail: illes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001143_acc.2013.6579933-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001143_acc.2013.6579933-Figure4-1.png",
        "caption": "Figure 4. Lumped System Modeling for Blade Flexibility",
        "texts": [
            " 2 2 2 2 2 2 sin( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 cos( ) sin( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 cos( ) cos( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 A a A A a A A a d g X u q w r v d R R M d g Y v r u p w d R R M d g Z w p v q u d R R M T \\ : : T I \\ : : T I \\ : : (1) 2 2 2 2 2 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 yy zz xz xx xx xx xx zz xx xz yy yy yy yyxx xz zz zz zz zz Id I I d L p q r p q r d I I I d I d I I I M q p r p q d I I I I yy Id I I d N r p q q r p d I I I d I \\ \\ : \\ : \\ \\ : \u00a7 \u00b7 \u00a7 \u00b7 \u00a8 \u00b8 \u00a8 \u00b8 \u00a9 \u00b9\u00a9 \u00b9 \u00a7 \u00b7 \u00a8 \u00b8\u00a8 \u00b8 \u00a9 \u00b9 \u00a7 \u00b7 \u00a7 \u00b7 \u00a8 \u00b8 \u00a8 \u00b8 \u00a9 \u00b9\u00a9 \u00b9 (2) The infinitesimal aerodynamic force and moment acting on a blade strip in lead-lagging and flapping frame (LLF) are 2 2 0 23 0 2 0 1 (3) 2 Pb P T P TLLF aero T T T P T UI U d F U U U dx U aR U U U U T G G T T \u00aa \u00ba \u00ab \u00bb \u00ab \u00bb \u00ab \u00bb \u00a7 \u00b7\u00ab \u00bb\u00a8 \u00b8\u00ab \u00bb\u00a9 \u00b9 \u00ab \u00bb \u00ab \u00bb \u00ab \u00bb \u00ac \u00bc J \u00a7 \u00b7\u00a7 \u00b7 \u00a8 \u00b8\u00a8 \u00b8\u00a8 \u00b8\u00a9 \u00b9\u00a9 \u00b9 0 ( ) ( ) T LLF aero III LLF aero IILLF aerod M xR d F xR d F\u00aa \u00ba \u00ac \u00bc (4) where P U and TU are perpendicular and tangential components of air velocity to the blade leading edge, T the blade pitch angle, xR the location of a point on the blade, 0a the blade lift curve slope, 0G , 2G parasite and induced drag coefficients. By integrating (3) and (4) along the blade span and including flapping and lead-lagging spring and damper moments, single blade equations are obtained [4]. For flexibility modeling, blades are divided into rigid pieces connected by flapwise bending springs and dampers (Fig. 4). The flapping angle of the (i+1) blade segment is 1 1 ( ) ( ) ( ) , 1, ..., 1 i i k k i nE \\ E \\ G \\ \u00a6 (5) where ( )kG \\ is the deflection angle of the k-th flapwise bending spring and ( )E \\ the root flapping angle. An equivalent energy approach [4] was used to find spring stiffnesses for n = number of blade segments = 3. Blade flapping, lead-lagging and flapwise bending motions are described, ignoring higher harmonic terms, by 0 1 0 1 ( ) cos( ) sin( ) 1 i i c i i d s \\ \\ \\4 4 4 4 4 (6) where ( / 2)( 1)i i\\ \\ S is the i-th blade azimuth angle, 4 is any of the three angles mentioned above, 0 4 , c4 , s4 , d4 are collective, two cyclic, and differential components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure8-1.png",
        "caption": "Fig. 8: Tube-to-plate welding case",
        "texts": [
            " These two chains constitute a closed kinematic chain. The plate-to-plate welding is a typical kind of tightly coupled cooperation multi-robot system, which requires the relative pose of end tools between dual robots remaining unchanged in the welding process, as Fig. 7 shows. Note that the signal t in the figure means tcp for simplification. The tube-to-plate and tube-to-tube welding cases are two typical kinds of loosely coupled robot coordination, which requires a gripping robot and a welding robot, as Fig. 8 and 9 shows. The gripping robot is treated as the maser and used to catch and transfer the workpiece into the expected pose with the ship welding constraint, meanwhile the welding robot do the welding job. According to (5), m sU is obtained by robot calibration. In order to obtain ( )s stU t , the master robot trajectory ( )m mtU t and the relative trajectory of the slave robot to the master robot tool ( )mt stU t should be calculated first. As the master robot, the end flange of the gripper robot is connected with the workpiece fixedly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002404_mmar.2016.7575286-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002404_mmar.2016.7575286-Figure10-1.png",
        "caption": "Fig. 10. The gripper frame x7y7z7 and object frame xobyobzob",
        "texts": [
            " For the robot station on which the studies were carried matrix Ts (see Fig.6) has the form (14). The matrix E describing the gripper frame x7y7z7 relative to the working link frame x6y6z6 has the form (15). It was assumed that the desired trajectory of the gripper relative to the object is a straight line coinciding with the axis zob. The gripper has to start straight motion from the point S and end at the point F above the object. It was assumed that the point S is at a height of 314 mm above the plane xobyob as in Fig.10. For such defined trajectory of the gripper we obtain the matrix obT7d(s) described by (16). 1 7 )300,0,0()( \u2212\u2212= \u03bcET sTranssd ob . (16)             \u2212 \u2212 = 1000 14100 0010 0001 \u03bcE . (17) At start point S the parameter s = 0, at the end point F s = smax = 300 mm. From Fig.10 results the matrix E\u03bc described by (17). Next the program dist0 calculated the matrix T6d(s) from equation (6). This matrix has been calculated at 301 points created from division the section SF = 300 mm on the 300 equal parts. From the matrix T6d(sl) at the points described by parameter s=sl (l=1,2,...,301, see (8)) the joint variables \u03981(sl)\u00f7\u03986(sl) of the Adept Six-300 robot were calculated. Figure 11 shows these variables. These joint variables were sent from the server to the robot controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002290_b15961-60-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002290_b15961-60-Figure2-1.png",
        "caption": "Figure 2. Tool for hot sheet metal forming.",
        "texts": [
            " Furthermore, it was assumed that by increasing the cooling rate, an improvement of the component\u2019s strength can be achieved. Subsequently, through the improved mechanical properties, a further reduction of wall thickness is getting possible. That results in a reduced need of raw material and therefore in resource savings. Due to the reduced cycle time, reduced energy requirements and possible material savings, energy savings of up to 10 per cent per produced part are expected to be possible. The setup of a hot forming tool (see Figure 2) is more complex than that of a conventional one. Mainly this is due to the fact that the cooling channels must be implemented into the punch and the die. The implementation of the channels is usually done by deep drilling or a segmentation of the tools. Due to complex geometry of the tools, the cooling system design is especially demanding for the tool manufacturer. The added complexity of cooling bores increases the expenses for hot forming tools. Current production effort is estimated with about one hour per meter borehole and a high consumption of resources (energy, drilling oil, compressed air, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.27-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.27-1.png",
        "caption": "Figure 2.27 Components of a glass micropipette.",
        "texts": [
            " These three types of microelectrodes are presented below: glass microelectrodes (glass micropipettes); metal microelectrodes; and microelectrodes fabricated using microelectronic technology. This electrode is fabricated from a glass capillary tube that is heated and pulled through its extremities forming a very narrow constriction in its center region ( 1 \u03bcm diameter). Then the tube is cut in the constriction forming two glass micropipettes. Each micropipette is filled with electrolyte (KCl 3 M) and receives a cap with a metal electrode rod or wire (Ag AgCl) coupled; this metal electrode is inserted in the electrolyte (Figure 2.27). The opening at the tip of the micropipette is the active region of the microelectrode and is large enough to allow ions exchange between the electrolyte and the outer environment (intra- or extracellular fluids). The glass capillary tube can be pulled and stretched manually (by very skilled people) but this is not recommended because of lack of reproducibility. Vertical and horizontal pullers with a platinum or nichrome heating element, preheat selectable times and one or two pull stages, allow to achieve good reproducibility, although setting them up correctly is time consuming"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003390_iccsp48568.2020.9182347-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003390_iccsp48568.2020.9182347-Figure1-1.png",
        "caption": "Fig. 1. Trajectories formed by different Bug Algorithm with reference to literature [20]",
        "texts": [
            " The sub goal points are the end points of the each black line. D. Algorithm Study and Analysis The fixed obstacles can categorise into two kinds: recognized and unrecognized. During offline planning few obstacles in advance are recognised by the planner, while few are not so to the planner and during navigation only it will be noticed by the sensor. When the mobile robot starts to traverse through the environment, it will start considering unrecognized stationary obstacles. The common view of avoiding an unknown static obstacle as per [20] is shown in Fig. 1. In this new approach the algorithm uses the direction and locations of the dynamic obstacle, whose path are indefinite to the robot at some stage. The algorithm is based on a strategy that every obstacle irrespective of stationary or moving is static at a particular point of time. If the position of an obstacle\u2019s varies time to time then we will say that the obstacle is dynamic and if the position of an obstacle remains same after some time interval t then the obstacle is static. Static obstacle can be avoided by the robot using Critical-PointBug algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000700_s11249-015-0616-0-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000700_s11249-015-0616-0-Figure1-1.png",
        "caption": "Fig. 1 Mechanism for controlling replenishment a schematic diagram and dimension of mechanism, b 3D view of the spherical roller-on-disk configuration with the slider",
        "texts": [
            " The spherical roller and disk are driven by servo motors controlled by the computer with the ability to change the required slide-to-roll ratio in a wide range. The slide-to-roll ratios are given as SRR \u00bc 2 \u00f0udisk uroller\u00de=\u00f0udisk \u00fe uroller\u00de and the entraining velocity as um \u00bc \u00f0udisk \u00fe uroller\u00de=2. The construction of the tribometer gives the ability of capturing interferometric images simultaneously with measuring the coefficient of friction. The schematic representation of test rig is described in an earlier work [21]. However, Fig. 1 shows dimensions and assembly of the slider and spherical roller-on-disk configuration. The elliptical contact created between the steel spherical roller and the glass disk has an ellipticity ratio of k = 1.84. In all experiments presented in this study, the contact has been loaded by 45 N resulting in a maximum Hertzian pressure of 0.445 GPa. A digital pipette has been used to supply measured volumes of lubricant on the surface of spherical roller close to its centerline. Then, the disk is mounted on the spherical roller, and the load is set. Next, the spherical roller and disk were run for 3 min under pure rolling to ensure a uniform distribution of lubricant on the tracks. Measurements were taken first for contact without slider under natural replenishment conditions; then, the slider is mounted as shown in Fig. 1 for measurements under channeled replenishment with the same amount of lubricant. All measurements were taken at ambient temperature of 22 C. The formation of the lubricating film is observed by colorimetric interferometry technique [22]. The principle of the method has been implemented in a comprehensive program that allows repeatedly measuring the film thickness under different experimental conditions. Lubricated contacts are observed using a microscope imaging system. This method utilizes interference of light at the interface of contacting bodies and the lubricating film",
            " The acquired signal from the torque sensor is processed by digital/analog card configured by software. The spherical roller is driven by a servo motor through torque sensor interfaced to computer. The sensor is well calibrated using static loads before all the tests. Also, compensation has been added to eliminate the effect of bearings on the spherical roller shaft. However, the sensor has absolute error of \u00b10.2 % according to the factory specifications. The mechanism required for achieving the channeling of lubricant is shown in Fig. 1. The implementation of this mechanism in spherical roller-on-disk test rig is quite easy using a slider (thin plastic sheet made of polyethylene). The thin plastic sheet is flexible with a tension force about 0.1 N to ensure a permanent contact with the surface of spherical roller with negligible effect on friction (friction between slider and surface of spherical roller is minor). In experiments presented in this paper, the slot of slider has a width of 0.8 mm and the width of slider head is 4",
            " The thickness of slider layer is 0.3 mm; indeed, the efficiency of scraping the track increases as the transverse width of the slider is larger. On the other hand, the profile of the contact between the slider and the spherical roller takes the shape of a chevron (the head of slider is chamfered upward with a small angle 12 on both sides) oriented transversely on the sliding direction resulting in scraping the lubricant without side leakage. The slider is attached firmly to a fixed substrate as shown in Fig. 1. The slot of slider is concentrated with the centerline of the spherical roller. Thus, the slider scrapes the lubricant on the surface of the spherical roller leading to a forced and channeled reflow through the slot toward the centerline. Two different base oils have been used in experiments to show the effect of viscosity on the mechanism of controlling replenishment and channeling lubricant. The first base oil (N2400) has a dynamic viscosity of 0.412 Pa.s at 40 C and a pressure\u2013viscosity coefficient of 35 GPa-1",
            " It is clear that the depleted track has a width of 583 lm which is very close to the width of Hertzian elliptical contact 596 lm. With such insufficient out-ofcontact replenishment, the resulting film thickness is about 150 nm, as shown in Fig. 2a. Under these starved conditions, the film thickness in the contact is formed essentially from the residual oil layer on the depleted track and by the effect of capillary forces in the vicinity of the contact [8, 11]. Introducing the mechanism of channeling replenishment to the spherical roller-on-disk configuration, as shown in Fig. 1, results in a significant flow modification and oil redistribution in the inlet of the contact. Figure 2b shows the track with channeled replenishment in relation to the size of Hertzian contact. The resulting track is fully flooded with fresh lubricant for the same volume of oil 10 ll. Although the slot of slider has a width of 800 lm, the scraped layer of oil has slightly less width (760 lm) due to the folding of slider a little on the surface of spherical roller to make sealed contact. However, the redistribution of oil on the surface of spherical roller leads to a fully flooded contact with central film thickness about 480 nm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000250_iros40897.2019.8968074-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000250_iros40897.2019.8968074-Figure3-1.png",
        "caption": "Fig. 3. Schematic figure of helical and parallel wire routing. Helical routing for outside and parallel routing for the inside of the PEFC tube to apply a uniformed distribution of voltage. One wire is used for helical routing, and a total of six wires are used for parallel wire routing: three pairs of wires with different lengths, respectively.",
        "texts": [
            " The proposed actuator has a dual cylindrical structure and is mainly composed of a thin McKibben muscle, a tube-shaped soft PEFC, and electrical wires for applying a voltage to the electrodes. The tube-shaped soft PEFC is inserted into the thin McKibben muscle, and its wall separates the inner space of the thin McKibben muscle into two rooms not to mix generated oxygen and hydrogen. Because of its high aspect ratio, it is difficult to apply a certain voltage to the whole body of the soft PEFC tube through normal methods via wire connections. Thus, we design a wire routing method named helical and parallel wire routing, illustrated in Fig. 3. A bare electrical wire just helically contacts the outside electrode to uniformly apply a voltage to the whole electrode without causing any effect of surface resistance. For the inside electrode, six bare electrical wires are inserted in parallel into the tube. These wires contact the inside electrode taking advantage of their restoring force. The optimal number of internal wires depends on parameters of the actuator like wire\u2019s stiffness, diameter, tube size, etc. On the one hand, a greater number of internal wires works better because it reduces the electrical impedance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001577_icuas.2013.6564739-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001577_icuas.2013.6564739-Figure1-1.png",
        "caption": "Fig. 1. AggieAir QuadrotorVTOL UAV platform",
        "texts": [
            " The designed FO[PI] controller is compared with an integer order PID (IOPID) controller and a MZNs PI controller, and the proposed fractional order controller outperforms both the MZNs PI controller and the integer order PID controller in terms of gain variations and minimizing disturbance. This paper is organized as follows: Section II introduces the AggieAir small VTOL UAV platforms, and describes the procedures of closed-loop identification for the pitch loop. Section III illustrates the design criterions and procedures to obtain the integer order MZNs PI controller, IOPID controller and FO[PI] controller. Simulation results are shown in Section IV. Section V concludes this paper. One of the AggieAir [16] small VTOL platforms is shown in Fig. 1 with specification given in Table I. AggieAir is a low cost Unmaned Aerial System (UAS) for personal remote sensing [17]. AggieAir uses the open- 978-1-4799-0817-2/13/$31.00 \u00a92013 IEEE 609 source Paparazzi autopilot [18], consumer grade electronics and sensors, while maintaining excellent flight characteristics and reliability [19]. The avionics of such small low cost UAVs consist of an Inertial Measurement Unit (IMU), which outputs attitude estimation, a GPS providing position information, pressure sensors for precise altitude estimation relative to a certain setpoint, a radio transmitter/receiver for telemetry remote control, and an autopilot unit which runs all control loops and stabilizes the attitude and altitude"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000129_ecce.2019.8911875-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000129_ecce.2019.8911875-Figure16-1.png",
        "caption": "Fig. 16. Mises stress distribution (8000 r/min)",
        "texts": [
            " The constraint condition is that the maximum Mises stress is less than 300 MPa at a over speed condition (8000 r/min) Fig. 14 shows the variation in the slit shape during the optimization. First, the slit inclines to the left side. Then, the width of the slit bottom at the right side is enlarged. Fig. 15 shows the variation in the torque with optimization iterations. The torque of the V-shape motor with the initial slit shown in Fig. 13 is considerably smaller than that of the V-shape motor without slit. On the other hand, it becomes larger by optimizing the slit as shown in Fig. 14. Fig. 16 shows the Mises stress distribution in the Vshape motor with optimized slit. It is confirmed that the maximum Mises stress is suppressed under the constraint condition. Fig. 17 shows the distributions of the total, PM, and armature reaction fluxes in the rotors with/without the optimized slit. It is observed that the optimized slit is nearly parallel to the total flux. It is also observed that the slit prevents the phase delay of the PM flux PM by the cross magnetization. On the other hand it prevents the armature reaction flux Iq produced by Iq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002827_j.optlaseng.2020.106065-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002827_j.optlaseng.2020.106065-Figure6-1.png",
        "caption": "Fig. 6. Schematic of the high-speed photographic experimental system.",
        "texts": [
            " 5) The surface of the powder particle, the side, and bottom surface of the calculation zone is defined as the wall boundary condition (WALL), while the upper surface of the calculation zone is defined as the pressure outlet boundary condition (PRESSURE_OUTLET). 6) A 3-D double precision solver is adopted to improve the accuracy and convergence of the calculations. .4. Model verification The validity of the model was confirmed using high-speed photoraphic experiments. Since the metal melt pool is non-transparent and ifficult to be observed, water is used as the liquid, and an experimenal measuring device shown in Fig. 6 was used to verify the interaction Fig. 4. Gambit model and meshing results: (a) gambit models on the entire calculation area; (b) particles and surrounding dynamic meshing results (partial magnification). Fig. 5. Micrographs showing typical unmelted powder particles (Ti-25V-15Cr-0.2Si flameresistant titanium alloy powder particles) just entering the interior of the deposited layer or partially entering the interior of the deposited layer: (a) a particle partially entering the interior of the deposited layer; (b) micrograph showing a particle completely entering the melt pool",
            " The particles re fed into the water surface vertically to facilitate observation of the ovement state of the particles after entering the liquid phase. Macrocopic high-speed photography was used to capture the change of the otion state of the particles passing through the gas-liquid interface, hich was then compared with the analytical results for the purpose f model verification. For the purpose of controlling movement and mount of the particles entering the water, a thin sheet with a small ole in the center was placed above the water tank (as shown in Fig. 6 ), hich ensures that only the particles passing through the small hole can nter the water. Also, the amount of the particles was reduced for easy dentification of particles. The final scheme of the verification experiment was determined ased on the influencing factors of the incident position of the particle, he position of the auxiliary light source, and the high-speed camera. uring the experiment, as shown in Fig. 7 (a), the powder feeding tube s installed near the wall of the water tank, to allow the water inlet of the owder particle is close to the wall of the water tank"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure18-1.png",
        "caption": "Figure 18. Velocity field contours for drum diameter 700 mm.",
        "texts": [],
        "surrounding_texts": [
            "Table 2 and the graph (Figure 19) show the calculations of the maximum speed of the velocity fields for a fixed drum diameter of 650 mm and different values of the drum rotation speeds. The diameter and rotation speed of the beater are equal to the initial values. [22] ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0002554_msnmc.2016.7783095-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002554_msnmc.2016.7783095-Figure1-1.png",
        "caption": "Fig. 1. Kineamtics diagram of the quad-rotor.",
        "texts": [
            " So, in order to synthesize the control law in this case, it is necessary to apply special algorithms proposed in [7]. Application of these algorithms allows the purposeful changing of the performance functional parameters in order to obtain desirable dynamic behavior of the flight control system. The results of the computer simulation demonstrate significant improvement of the LQR-based quad-rotor flight control system performance in comparison with results obtained in [4]. II. EQUATION OF MOTION The quad-rotor scheme is represented in Fig. 1 (corresponds to Fig. 3 in [3]). Let  x y z   is the radiusvector of the quad-rotor center of mass in some inertial frame {X, Y, Z}, , ,   are the yaw, pitch and roll angles respectively, if is the lifting force produced by the i th motor iM ( 1,4)i  . Here and further prime denotes transposing. In accordance with [3], [4], the motion of this system is described by the following system of equations: 2 2 sin , d xm u dt   (1) 2016 4th International Conference on Methods and Systems of Navigation and Motion Control (MSNMC) Proceedings 13 2 2 cos sin ,d ym u dt    (2) 2 2 cos cos ,d zm u mg dt     (3) 2 2 ,d dt     (4) 2 2 ,d dt     (5) 2 2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001025_2015-01-1477-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001025_2015-01-1477-Figure1-1.png",
        "caption": "Figure 1. Graphical representation of intended steering-induced rollover test layout.",
        "texts": [
            " Both vehicles were accelerated to the target test speed with a cable-guided tow system, released on a paved tow lane, and an automated steering controller input a right steer followed by a rapid left steer. The right steer caused the vehicle to travel off the paved tow lane and onto a prepared off-road soil surface, where a rapid left steer was input that caused the vehicle to yaw counter-clockwise and enter a passenger-side leading yaw orientation. The furrowing of the passenger side tires induced an off-road soil-tripped rollover, as illustrated in Figure 1. The testing was documented with an extensive array of instrumentation on the vehicle and the ATDs. Real-time highdefinition (HD) and high-speed video was recorded both onboard and offboard, which captured the handling maneuver, the trip phase, and the rollover event. One additional pickup test and several more sedan tests were attempted, but the vehicles did not roll over; instead they yawed and slid to rest. An additional sedan test had a steering controller malfunction that resulted in the vehicle rolling over after traversing an adjacent ditch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000041_s11668-019-00763-2-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000041_s11668-019-00763-2-Figure2-1.png",
        "caption": "Fig. 2 The lower arm of MF285 tractor: a actual and b the simulated one in SolidWorks",
        "texts": [],
        "surrounding_texts": [
            "In the present study, the dimensions of the lower arms of 399 and 285 Massey Ferguson tractors were measured. These were used to simulate a three-dimensional mechanism of lower arms in SolidWorks 2013, a commercial software. The real and schematic lower arms of MF399 and MF285 tractors are shown in Figs. 1 and 2, respectively. Then, the CAD simulation files in order to build the FEM were imported into ANSYS V15 Software. Mechanical properties of steel (St 37) which was used to build the lower arms are presented in Table 1. Static Analysis The finite element models were meshed using SOLID187 that is a three-dimensional solid element and has 8 nodes with three degrees of freedom for each node. The models consisted of 614 elements and 16,318 nodes for the lower arm of MF399 tractor and 530 elements and 99,089 nodes for the lower arm of MF285 tractor (Fig. 3). Meshing of models was performed by the free method. The boundary and displacement conditions were applied to arm junction points to the tractors. The values of loads applied to the lower arms of MF399 and MF285 tractors were obtained from field testing under different working conditions. The tensile force testing was performed in order to determine maximum value of forces on the lower arms and its direction while working with chisel, furrower equipment and drill planter. The lower arms supply draft forces which are applied in parallel with the forward direction of tractor. The tensile load of chisel plow and furrower equipment during tillage operation and drill planter while planting were measured at the maximum depth of working (Table 2) and were considered as input loads for analyzing of the model in ANSYS software. In static analysis, force was applied to the lower arms and after analyzing the model, the equivalent tension stress on the lower arms was calculated based on Von-Mises theory. To avoid of the parts failure, the maximum stress must not exceed of the material yielding strength. Under static loading, the factor of safety (FS) is obtained from yield strength (ryp) divided by maximum working or allowable stress (rall) applied on the lower links (Eq 1). In a proper design, the FS must be greater than one, indicating that the structure can work and remain healthy in uncertain loading conditions [21]. F:S: \u00bc ry rall \u00f0Eq 1\u00de Modal Analysis Modal analysis is necessary in the structural and industrial components designs that can resonate. Small loads at resonance frequency can result in induced deformation and damage in the structures. Therefore, mechanical parts must be designed as far as possible away from the resonance Table 1 Mechanical properties of steel (St 37) Density, q (kg/ m3) Modulus of elasticity, E (GPa) Poisson\u2019s ratio, t Yielding strength, ryp (MPa) Ultimate tensile strength, rut (MPa) 7860 200 0.3 198 235 Table 2 Equipment properties and the maximum loads applied to lower arms Implements Number of unit Maximum depth (cm) Tensile load (KN) Furrower 4 24 8.64 Chisel plow (sweep) 9 24 9.71 Drill planter 6 12 14.22 frequency range. Fluctuation in natural frequency of the structure increases the vibration amplitude and results in the failure and fracture of the component. Modal analysis is used to determine the value and mode shapes of the natural frequencies [22]. The natural frequency of the structure depends on the shape, material and supports of the structure. However, the amount and type of loads can affect the natural frequency. In this study, the SOLID 182 and 185 elements were used for modal analysis of the parts. To determine the characterization of materials including modulus of elasticity, Poisson\u2019s ratio and density of steel (St 37) were used (Table 1). The models were meshed, boundary conditions were applied in the lower arms connecting to the tractor, and desired nodes were binding in all directions. Then, the modal analysis of the lower arms of MF399 and MF285 tractors was performed with regard to the first 5 natural frequencies in range of 0\u20135000 Hz. Fatigue Analysis The effect of cyclic loads applied on structures creates cracks and finally results in part fracture, while the magnitude of reversing stress is lower than the yield stress of the structure. This phenomenon is called fatigue due to applying cyclic loads. In ANSYS software to perform a fatigue analysis under intermittent loads, firstly the exerted stresses in structures under cyclic loads must be determined. So, before any fatigue analysis, statistical analysis should be performed. Then according to stress contours, the critical nodes with maximum stresses must be detected and after fatigue analysis can be surveyed on the critical nodes. In order to obtain the endurance stress (re) of the steel material in the lower arms, fatigue analysis was performed on the model by applying 1.5 million loading cycles. For this purpose, the SOLID 185 element was used. It is a three-dimensional and 4-node element, and each node has three degrees of freedom. Characterization of materials, mesh and applying boundary conditions was conducted the same as those considered for static analysis. The Soderberg equation (Eq 2) was used to calculate factor of safety (FS) in fatigue analysis [23]. 1 F:S: \u00bc rave ry \u00fe K rr re \u00f0Eq 2\u00de where rave is average stress, rr is reversing stress, and K is geometric stress concentration factor. In the above equation, rave and rr were obtained by Eqs 3 and 4. In these equations, rmax was the maximum stress and rmin was considered as zero. rave \u00bc rmax \u00fe rmin 2 \u00f0Eq 3\u00de rr \u00bc rmax rmin 2 \u00f0Eq 4\u00de"
        ]
    },
    {
        "image_filename": "designv11_30_0002962_1.5133369-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002962_1.5133369-Figure2-1.png",
        "caption": "FIG. 2. The schematic representation of the strut. 1, universal Joint; 2, linear actuator; 3, transmission nut; 4-I, interior buffer material; 4-II, external buffer material; 5, inner tube; 6, outer tube; 7, stretch tube; 8, spherical joint.",
        "texts": [
            " However, in order to reduce the stress and stretch from the primary strut, two cushioning materials were installed on the secondary strut. In each parallel leg mechanism, linear actuators were installed on the primary strut and secondary strut. Ensure that the parallel mechanism can complete the drive operation. Because the lander has a higher center of gravity (the initial height depends on the length of the linear actuator) and a large walking inertia, there is a high requirement for the stability of the lander\u2019s motion after landing.34 In Fig. 2, its soft-landing is like that of a conventional lander. In order to balance buffering and walking, it is necessary to change the installation position and structure of the buffer material. The ring structure buffer material is designed to avoid the interference in transmission mechanism from the compressed buffer material. During the soft landing, only the primary strut bears the pressure. The inner tube transmits pressure to the material to buffer the impact. As a result of the way it lands, the support pillars bear a large magnitude of pressure and tension"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure5-1.png",
        "caption": "Fig. 5. Block diagram (a); mechanism diagram (b) of the 3-RRPRR serial chains of TPM (R: Revolute joints, P: Prismatic joints, E: Planar joint (3R), EE: End-effector).",
        "texts": [
            " To eliminate the rotational movement of the piston rod of the hydraulic actuator around its axis, an additional rotation lockout was used. Fig. 3 shows the rotation lockout and provides a description of the individual elements. Pure translational motion of the mobile platform of the PM is possible after using a specially designed planar joint (Fig. 4). Here, the hydraulic actuator piston rods are mounted on pins in the indi- vidual rotary joints. The moving platform is mounted on the main pin, which is the end-effector (EE). The TPM kinematic model with three 3-RRPRR serial chains is shown in Fig. 5. The TPM design consists of a fixed platform, a mobile platform, twelve passive revolute joints (R), including a planar joint (E) and three active prismatic joints (P). The Chebychev-Gr\u00fcbler-Kutzbach (CGK) formula was used to determine the mobility (m = 3) of the hydraulic TPM [23]. Experimental studies of the 3-RRPRR TPM with three IEHSDs revealed unexpected additional degrees of freedom. These come about when the hydraulic actuator piston rods are not blocked from rotating. In such situations, the kinematic mechanism of the TPM, instead of being rigid, behaves as if it had additional degrees of freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002409_cca.2016.7587991-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002409_cca.2016.7587991-Figure5-1.png",
        "caption": "Fig. 5: Region of interest extraction procedure. First, the shaft is removed from the image (a). Afterwards the outer-raceway is detected (b). Finally, the background is removed (c) and the left part of the image is removed (d).",
        "texts": [
            " The goal in the ROI extraction step is to segment the right side of the bearing housing as the plexiglass part cannot be used for feature extraction as plexiglass is not used in industrial machines. The plexiglass part, in this set-up, is used to visually monitor the oil-level and create ground truth labels for the different test runs. Furthermore, only the cover should be extracted and the background should be removed as the cover is closest to the bearing. To do this ROI segmentation, several steps are required. First, the shaft has to be detected. As the shaft is a circle in the image, this is done using the circle Hough transform [15]. The result of this can be seen in Fig. 5a. Using this circle a binary matrix, i.e. a mask, is created according to equation (2) where M1 is the binary matrix and i and j its indices, i1c and j1c are the coordinates of the center of the circle and r1 its radius. In this binary mask all pixels within the circle will be equal to 0 and the pixels outside the circle equal to 1. Afterwards this mask is applied on the IRT frames: I = I M1. M1ij = { 0, if (i\u2212 i1c)2 + (j \u2212 j1c)2 < r21 1, otherwise. (2) Now that the center is removed, in the next step the circle Hough transform is applied again. This time to detect the outer-raceway of the bearing. The result can be seen in Fig. 5b. The goal is now to remove the background of the image. To do this, again a mask is created as in equation 3. This time, the pixels within the circle are equal to 1 and those outside the circle equal to 0. Afterwards, the mask is applied on the IRT frames: I = I M2. This operation results in an isolated ROI as can be seen in Fig 5c. From this ROI only the right-hand side is kept as this side made out of stainless steel. The final ROI can be seen in Fig 5d. M2ij = { 0, if (i\u2212 i2c)2 + (j \u2212 j2c)2 > r22 1, otherwise. (3) The ROI extractions step is applied on every frame. Afterwards, per frame several features are extracted: \u2022 Tmax, which is the maximum temperature \u2022 Tmin, which is the minimum temperature \u2022 \u2206Tmax\u2212min = Tmax \u2212 Tmin \u2022 Ttop = temperature at the top of the ROI \u2022 Tbottom = temperature at the bottom of the ROI \u2022 \u2206Ttop\u2212bottom = Ttop \u2212 Tbottom These features are chosen as we expect the bottom of the cover to have a different temperature compared to the top of the cover"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002524_s00419-016-1210-0-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002524_s00419-016-1210-0-Figure1-1.png",
        "caption": "Fig. 1 Beam element: local coordinates in the two bending planes and shear effect on a small section [14]",
        "texts": [
            " The definition of the vector differential equation for the complete rotor-dynamic system is conducted stepwise by assembling the system matrices for the shaft, the wheels and bearings, respectively [9,14,15,30,32]. The shaft is studied section-wise by utilizing the displacement and rotation of each element with the help of the Timoshenko beam theory [33]. Here, both the notation and modelling approach are adapted as formulated in [14] and are presented in compact form throughout the rest of the section. The beam element matrices are initially obtained on the basis of one bending plane by accounting both shear effects and rotary inertia within the definition of the element displacement as depicted in Fig. 1 and (1)\u2013(7): ue,i (\u03be, t) = [ N1(\u03be) N2(\u03be) N3(\u03be) N4(\u03be) ] [ ue,i (t) \u03b8e,i (t) ue,i+1(t) \u03b8e,i+1(t) ]T (1) \u03b8e,i (\u03be, t) = \u2202ue,i (\u03be, t) \u2202\u03be + \u03b2e,i (\u03be, t) (2) N1(\u03be) = 1 1 + \u03a6e ( 1 + \u03a6e \u2212 \u03a6e \u03be le \u2212 3 \u03be2 l2e + 2 \u03be3 l3e ) (3) N2(\u03be) = le 1 + \u03a6e ( (2 + \u03a6e) \u03be 2le \u2212 (4 + \u03a6e) \u03be2 2l2e + \u03be3 l3e ) (4) N3(\u03be) = 1 1 + \u03a6e ( \u03a6e \u03be le + 3 \u03be2 l2e \u2212 2 \u03be3 l3e ) (5) N4(\u03be) = le 1 + \u03a6e ( \u2212\u03a6e\u03be 2le \u2212 (2 \u2212 \u03a6e) \u03be2 2l2e + \u03be3 l3e ) (6) \u03b8e,i (t) = \u2202ue,i \u2202\u03be |\u03be=0, \u03b8e,i+1(t) = \u2202ue,i \u2202\u03be |\u03be=le + \u03b2e,i (le, t), \u03a6e = 12Ee Ie keGe Ael2e (7) with ue,i (\u03be, t) denoting the approximated deflection within the i-th beam element on the basis of the Ni (\u03be) shape functions, \u03b8e,i (\u03be, t) the angle of the beam cross section and \u03b2e,i (\u03be, t) the shear effect. Here, only solid shafts are investigated with the shear constant ke = 6(1+\u03bde) 2 (7+12\u03bde+4\u03bd2e ) [14], the shear modulus Ge = Ee 2(1+\u03bde) and the Poisson\u2019s ratio \u03bde. Thereafter, the second-bending plane (Fig. 1) is considered for the generalized formulation of the i-th element\u2019s lateral translation and rotation in the arbitrary position \u03be along the element\u2019s length (8)\u2013(9). It is conducted based on the generalized local coordinated vector Qe,i = [ ui , vi , \u03b8i , \u03c8i , ui+1, vi+1, \u03b8i+1, \u03c8i+1 ]T [14,32,33], i.e. \u03b8e,i (\u03be, t) = due,i d\u03be , \u03c8e,i (\u03be, t) = \u2212dve,i d\u03be (8) \u23a1 \u23a2 \u23a3 ue,i (\u03be, t) ve,i (\u03be, t) \u03b8e,i (\u03be, t) \u03c8e,i (\u03be, t) \u23a4 \u23a5 \u23a6 \ufe38 \ufe37\ufe37 \ufe38 qe,i = \u23a1 \u23a2 \u23a3 N1 0 0 N2 N3 0 0 N4 0 N1 N2 0 0 N3 N4 0 N \u2032 1 0 0 N \u2032 2 N \u2032 3 0 0 N \u2032 4 0 \u2212N1 N \u2032 2 0 0 \u2212N \u2032 3 N \u2032 4 0 \u23a4 \u23a5 \u23a6 Qe,i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.67-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.67-1.png",
        "caption": "Figure 3.67 Lead screw cross-section",
        "texts": [],
        "surrounding_texts": [
            "These components basically operate in the same manner. A nut in contact with an externally threaded rod, the screw, moves laterally as the screw is rotated. The major difference between a lead screw and a ball screw is the nature of the nut. In a lead screw an internally threaded nut mates with the screw. Contact between the two members consists of sliding friction. In a ball screw, the sliding friction of the lead screw nut is replaced with the rolling friction of ball bearings which circulate around the screw threads as the screw rotates. See Figures 3.67 and 3.68. In the majority of applications, the assembly is used to convert rotary motion into linear motion by driving the screw with a motor (DC, AC or stepper) and having the load, connected to the nut, move laterally. There are also assemblies in which the nut is mounted internal to the rotor of a motor and the screw moves laterally as the rotor/nut rotates. In addition, due to its low friction and high efficiency, it is possible to move the nut of a ball screw assembly laterally, causing the screw to rotate in a fixed mount. The following specifications represent a summary of data compiled from a number of manufacturers. 120 Electromechanical Motion Systems: Design and Simulation"
        ]
    },
    {
        "image_filename": "designv11_30_0003802_vppc49601.2020.9330853-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003802_vppc49601.2020.9330853-Figure2-1.png",
        "caption": "Fig. 2. Volume reduction of SiC inverter compared with IGBT inverter",
        "texts": [],
        "surrounding_texts": [
            "vehicles In this paper, a full-SiC traction inverter for metro vehicle is in use for studying, and the basic information of the inverter system is shown in Figure 1 ,2. The parameters also are as Table I showed. As mentioned above, the new inverter is nearly 30% smaller in terms of both weight and volume than IGBT inverter under the same driving condition of 1 inverter driving 4 motors. 978-1-7281-8959-8/20/$31.00 \u00a92020 IEEE 20 20 I E E E V eh ic le P ow er a nd P ro pu ls io n C on fe re nc e (V PP C ) | 9 78 -1 -7 28 1- 89 59 -8 /2 0/ $3 1. 00 \u00a9 20 20 I E E E | D O I: 1 0. 11 09 /V PP C 49 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 22,2021 at 20:41:28 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_30_0002968_lra.2020.2986746-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002968_lra.2020.2986746-Figure6-1.png",
        "caption": "Fig. 6. (a) When the gripper is inflated past ambient, it risks buckling which complicates data processing. (b)-(d) Top and isometric views of the gripper as it measures an extruded (yellow) shape with convex and concave features. The red arrows indicate direction of motion.",
        "texts": [
            " Note, that the syringe pump was used for experimental ease, but the inflation could also have been done by hand. The total fabrication time for the gripper array was no more than 5 hours from start to completion. We use a syringe pump to actuate the pneumatic gripper at the speed of 50 mL / s, causing a change in stretch at the same order of that used in the characterization of the single sensor (Sec. II-C). Note that using inflation alone would cause the material to buckle inwards, which would complicate sensor response and consequently the shape reconstruction (Fig. 6a). Instead, we avoid sensor buckling by only using the membrane under stretch. To sense the geometry of an encompassed object, we therefore conduct the following steps also shown in Fig. 6(b-d): (1) we deflate the gripper for 5 s to reach negative pressure, dilating the center of the gripper; (2) we hold a negative pressure for 5 s, during which we insert the object to be measured; (3) we Authorized licensed use limited to: University of Exeter. Downloaded on May 07,2020 at 19:52:16 UTC from IEEE Xplore. Restrictions apply. then inflate for 6 seconds, allowing the gripper to encompass the object fully. Pressure inside the gripper goes from ambient to negative (deflation and hold), to ambient and then positive pressure (inflation) subsequently"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002635_ecce.2016.7854747-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002635_ecce.2016.7854747-Figure1-1.png",
        "caption": "Fig. 1. The proposed 6/7-pole SCPMM.",
        "texts": [
            " In this paper, the topology and operating principle of the machine will be introduced, respectively. Afterwards, the available slot/pole combinations are analyzed. Furthermore, the electromagnetic performances of the proposed machines having different slot/pole combinations are investigated and compared. Finally, the prototype is manufactured and tested to experimentally validate the finiteelement (FE) analysis. II. MACHINE TOPOLOGY The topology of the SCPMM with 6/7 stator slot/rotor pole is shown in Fig. 1(a), in which the AlNiCo PMs are alternately 978-1-5090-0737-0/16/$31.00 \u00a92016 IEEE arranged between the adjacent stator poles. The consequentpole PMs are with the same magnetization directions. The magnetizing coils are accommodated in the slot between the stator pole and PM. Similarly to other stator excited machines [12], the rotor of the new machine is simply composed of salient iron poles, which provides excellent structural simplicity, robustness and ease of manufacture. Besides, the ferromagnetic poles provide effective circulating path for the armature fields"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003346_j.matpr.2020.07.346-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003346_j.matpr.2020.07.346-Figure1-1.png",
        "caption": "Fig. 1. Two Stages of Leaf Spring Design.",
        "texts": [
            " The introduction of composite materials made it possible to reduce the weight of the leaf spring without reduction of load carrying capacity and stiffness due to more elastic strain energy storage capacity and High strength to weight ratio. Please cite this article as: K. Krishnamurthy, P. Ravichandran, A. Shahid Naufal e rials, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.07.34 Initially the leaf spring has to be designed in the SolidWorks software as a 3D model. The type of spring we designed here is a general shape of multi-leaf spring for light trucks. Fig. 1 shows a general shape of the popular spring type, particularly for the light trucks which consist of 2 stages. The first stage having 3 leaves from leaf 1 to leaf 3, whereas the second stage, which is called helper spring, having 2 leaves. The character of this model is variable or progressive rate multi-leaf spring that ride qualities can be regulated over a wide load range. Once the designing phase is over in the software, then to process the analysis we need to get to ANSYS software. To convert the SolidWorks file into ANSYS we need save the same file in the format of IGES (International Graphic Exchange System) file"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003850_10_2015_324-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003850_10_2015_324-Figure3-1.png",
        "caption": "Fig. 3 Three-dimensional modeling performed by computer-aided design (left) and photos (right) of the Lumisens IV instrument (a global view; b inside of the measuring chamber). Reprinted with permission from Jouanneau et al. [8], copyright 2012 ACS",
        "texts": [
            " The bioluminescence images produced by the immobilized bacteria in agarose hydrogel arrays were recorded and quantified with a CCD cooled camera in a dark chamber, followed by image processing via customized software. This device can be used for 7 days in laboratory conditions, although its application in field has been limited to only 2 days because of oxygen shortage. An updated prototype built on the basis of freeze-dried bacteria shows confident bioluminescent detection (with 3 % reproducibility) for 10 days in both laboratory conditions and environmental conditions [8]. An overview of the latest CCD-based bioluminescent bacteria biosensor arrays is shown in Fig. 3. Because of the addressable configuration of the Lumisens IV instrument working in online mode, the \u201cfingerprint\u201d of a given pollutant or the total toxicity of various pollutants can be simultaneously investigated by using different populations of bacterial cells placed in a specific array positions. Considering the non-specificity of promoter genes in the engineered bacterial cells, the use of multi-strains is a promising alternative to rapidly improve the selectivity of biosensors in a real time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000918_icems.2013.6754383-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000918_icems.2013.6754383-Figure2-1.png",
        "caption": "Fig. 2. Structure of third layer.",
        "texts": [
            " The permanent magnet, NMF-12E, is the ferrite magnet, and the stator has distributed windings. Size and rated values of the motor proposed are the same as a motor with bar rare-earth magnets for 15 kW. Fig. 1 shows the motor proposed. This motor produces higher reluctance torque by arranging magnets along the magnetic flux flow [4]. It consists of three magnet layers [5], and there are ribs in the third layer. The width of the rib is 0.5 mm. III. RESULTS OF ANALYSIS The motor proposed with ribs and the motor Type A with no ribs in the third layer are shown in Fig. 2(a) and (b), respectively. The characteristics of the motor proposed are compared with Type A to demonstrate the effect of the ribs. 978-1-4799-1447-0/13/$31.00 \u00a92013 IEEE A. Torque Characteristic Fig. 3 and 4 show the relation between the average torque and the phase angle \u03b2 of the motor proposed and Type A, respectively. The maximum total torque is obtained when \u03b2 is 50 degrees because the reluctance torque is high as shown in these figures. The maximum torque of motor proposed is 56.1 N\u30fbm and that of Type A is 57",
            "4 kW as shown in Fig. 7. The output of the motor proposed decreases due to the torque reduction. Fig. 8 shows the relation between the phase angle \u03b2 and the rotor speed. As the rotor speed becomes higher, the phase angle \u03b2 becomes larger to maintain a line voltage of 400 Vrms as shown in this figure. There is almost no difference of \u03b2 between the two motors. C. Mechanical Strength of Rotor Fig. 9 shows the stress distribution at 10,0 in the third layer are made along the normal d motor\u2019s axis as shown in Fig. 2(a). The improves the mechanical strength because th directly press the ribs which are made a direction. The stress at the edge of the magn shown in Fig. 9. The yield stress of iron core is 275 MPa, a maximum stress is calculated by (1). a y sF \u03c3 \u03c3 = where sF safety factor y\u03c3 yield stress a\u03c3 allowable maximum stress When a safety factor is 1.5, the allowable is 183.3 MPa. The maximum speed is d maximum stress in the rotor reaches 183.3 M results of analysis, the maximum speed of the is 10,900 rpm and that of Type A is 8,300 rpm factor is considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001228_j.apm.2013.09.004-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001228_j.apm.2013.09.004-Figure2-1.png",
        "caption": "Fig. 2. Inverted pendulum system.",
        "texts": [
            " Assuming that the initial condition of the state is x0 \u00bc 0:2 0 0:2\u00bd T , the closed-loop state responses are depicted in Fig. 1. From the figure, we can see that the state variables asymptotically converge to zero under the effects of time-delay, packet dropout and quantization error, which shows the effectiveness of the controller design method in Theorem 2. Example 2. Suppose the simplified system in (26) is an inverted pendulum system with delayed control input. The inverted pendulum on a cart is shown in Fig. 2. In this system, a pendulum is attached to the side of a cart by means of a pivot which allows the pendulum to swing in xy-plane. A force u is applied to the cart in the x direction, with the purpose of keeping the pendulum balanced upright. x is the displacement of the center of mass of the cart from the origin O; h is the angle of the pendulum from the top vertical; M and m are the masses of the cart and the pendulum, respectively; l is the distance from the pivot to the center of mass of the pendulum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003573_0142331220937895-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003573_0142331220937895-Figure1-1.png",
        "caption": "Figure 1. Spacecraft model using multi-mass modeling (Sidi, 1997).",
        "texts": [
            " Applying a torque T on its axis of rotation leads to a motion u with rigid body and a deformation d from the rigid body. Using Lagrange\u2019s method, the motion equations can be derived as (1) I + 2mL2 \u20acu+ 2mL\u20acd =T m\u20acd +KM _d +KEd = mL\u20acu \u00f01\u00de Here, I is the moment of inertia of the rigid body, T is the applied torque on axis of rotation while KM and KE are mechanical dissipative constant and potential energy constant, respectively. Using the mass-spring concept and multi-mass modeling, the spacecraft model in the body-fixed coordinate can be derived, which is shown in Figure 1. In Figure 1, XB,YB,ZB are the body reference frame axes of the spacecraft. With such a concept, the equation of motion can be written as (2) M \u20acd +D _d+Ksd= \u20acuMx \u00f02\u00de where M is the mass matrix, Ks is the stiffness matrix, x is the distance and D is added for damping factors. To have a better understanding of a flexible satellite, a simple spacecraft model as two rigid bodies connected at a point is depicted in Figure 2. According to Figure 2, X, Y, Z describe the orientation of the spacecraft in the body-fixed reference frame while Xp, Yp, Zp show the reference frame of the appendages"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002155_icciautom.2016.7483170-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002155_icciautom.2016.7483170-Figure1-1.png",
        "caption": "Fig. 1. Quad rotor dynamics[12]",
        "texts": [
            " In this paper we proposed fuzzy controllers with optimized membership functions based on Particle Swarm Optimization (PSO) and Genetic Algorithm (GA) to control both transitional and angular movements of a quad rotor. In the following sections a brief discussion of the quad rotor configuration and dynamic model is presented. Then, the flight control algorithm is discussed. After that, PSO and GA algorithm are concisely introduced. Finally, we show the simulation results to illustrate the effectiveness of the proposed methods on the fuzzy controller\u2019s performance. According to Fig.1. the front and rear rotors rotate clockwise while the right and left rotors rotate counterclockwise. Increasing (decreasing) the speed of all rotors with the same amount generates the vertical motion. A roll motion can be obtained by increasing (decreasing) the left rotor\u2019s speed while decreasing (increasing) the speed of the right rotor. Similarly, a pitch motion is controlled by an increase (decrease) in the speed of the rear rotor while decreasing (increasing) the front rotor\u2019s speed. Yaw angle is achieved by increasing the speeds of clockwise pair and decreasing the speeds of counter-clockwise pair simultaneously"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000550_asjc.930-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000550_asjc.930-Figure2-1.png",
        "caption": "Fig. 2. Multi-agent system to be studied.",
        "texts": [
            " That is, G(x) is the undirected graph with the node set {1, 2,\u2026 , n} and the edge set{ (i, j) \u2208 {1, 2,\u2026 , n}2 |||Vi(S, x) \u2229 Vj(S, x) \u2260 \u2205 } . On the graph G(x), if node j is connected to node i, then node j is called the neighbor of node i. Next, let cent(S, f ) \u2236= \u222b S sf (s)ds \u222b S f (s)ds where S \u2282 R2 is a bounded and nonzero-measure set and f \u2236 R2 \u2192 R0+ is a function. This represents the weighted centroid of the set S. Finally, for the vectors x1, x2,\u2026 , xn \u2208 R2 and the set I \u2236= {i1, i2,\u2026 , im} \u2286 {1, 2,\u2026 , n}, let [xi]i\u2208I \u2236= [ x\u22a4 i1 x\u22a4 i2 \u00b7 \u00b7 \u00b7 x\u22a4 im ]\u22a4 \u2208 R2m. Consider the multi-agent system in Fig. 2. This is composed of n agents, and the dynamics of agent i (i \u2208 {1, 2,\u2026 , n}) is given by x\u0307i(t) = ui(t) (1) where xi(t) \u2208 R2 and ui(t) \u2208 R2 are the position and the input, respectively. The collective position of the agents is denoted by x(t) \u2208 R2n, i.e., x(t) \u2236= [ x\u22a4 1 (t) x\u22a4 2 (t) \u00b7 \u00b7 \u00b7 x\u22a4 n (t) ]\u22a4 , and x(t) is called the formation at time t. The initial formation is given as x(0) \u2236= x0 \u2208 R2n. Furthermore, if all the agents exist in a bounded convex set S \u2282 R2, we use Vi(S, x(t)) to express the Voronoi cell for xi(t) (for agent i) and use G(x(t)) to represent the corresponding Delaunay graph",
            " Here, \ud835\udf09i(t) \u2208 Rm is the state, [ xj(t) ] j\u2208Ni(t) \u2208 R 2|Ni(t)| is the input, ui(t) \u2208 R2 is the output, and f \u2236 Rm\u00d7R2|Ni(t)|\u00d7R0+ \u2192 Rm and g \u2236 Rm\u00d7R2|Ni(t)|\u00d7R0+ \u2192 R2 are functions. The input is composed of the positions of itself and its neighbors, and in this sense Ki is a distributed controller. The functions f and g and the initial state \ud835\udf09i(0) are assumed to be the same for all the agents. For simplicity, we further assume \ud835\udf09i(0) = 0. (3) Next, we formulate the mass game problem. For the multi-agent system in Fig. 2, suppose that a set S \u2282 R2, a convex set F \u2282 S, and a grayscale image r \u2236 F \u2192 [0, 1] are given. The set S is the field where the agents exist and F is the field where the agents organize themselves into a formation displaying the grayscale image r. The complement S\u29f5F is the space for evacuating unnecessary agents in this purpose. The grayscale image r is defined as the function expressing the pixel values on F, and the values r(p) = 0, r(p) = 1, and r(p) \u2208 (0, 1) correspond to black, white, and gray, respectively",
            " Based on these, our controllers are developed. 3.1 Elemental techniques for mass games 3.1.1 Coverage control The coverage means steering agents arbitrarily placed in an environment to the locations such that the sizes of the agents\u2019 occupied areas are equal in a certain sense. It is one of the fundamental coordination tasks in various multi-agent problems including mobile sensor networks. A theoretical framework for distributed coverage control has been developed in [16], and is summarized as follows. Consider the multi-agent system in Fig. 2. For \u00a9 2014 Chinese Automatic Control Society and Wiley Publishing Asia Pty Ltd the formation x \u2236= [ x\u22a4 1 x\u22a4 2 \u00b7 \u00b7 \u00b7 x\u22a4 n ]\u22a4 \u2208 R2n, we use the performance index J(x) \u2236= \u222b S min i\u2208{1,2,\u2026,n} \u2016p \u2212 xi\u20162\ud835\udf19(p) dp (4) where S is a bounded set representing the coverage field and \ud835\udf19 \u2236 S \u2192 R0+ is an integrable function which corresponds to the weighting function quantifying the relative importance of each point in S. By noting that mini\u2208{1,2,\u2026,n} \u2016p \u2212 xi\u20162 represents the distance between the point p and the nearest agent\u2019s position, the formation x minimizing J(x) under \ud835\udf19(p) \u2261 1 is a configuration that xi exists near at any point in S",
            " Second, the proposed controllers are distributed. In fact, Ki1 and Ki3 are controllers depending on the position of agent i itself, and Ki2 is also a distributed controller because the Voronoi cell Vi (F, x\u0304(t)) can be determined by the positions of agent i and its neighbors in the field F. In addition, the switching rule depends on the positions of itself and its neighbors in F since \ud835\udefe (x(t), a) is a function of Vi (F, x\u0304(t)). For the proposed controllers, the following result is obtained. Theorem 1. For the multi-agent system in Fig. 2, suppose that S \u2282 R2, F \u2282 S, and r \u2236 F \u2192 [0, 1] are given. Let K1,K2,\u2026 ,Kn be given by (2), (6), and (7). Let also x\u2217 \u2208 Sn be a formation such that the positions of the agents in F are a local minimum point of J where S is replaced by F. If x0 \u2208 (F \u29f5 cent(F, 1))n and c > 1 + k min (p1,p2)\u2208\ud835\udf15S\u00d7\ud835\udf15F \u2016p1 \u2212 p2\u2016 , (11) then x(t) \u2208 Sn for every t \u2208 R0+ and lim t\u2192\u221e x(t) = x\u2217. (12) Proof. The following three facts prove the theorem. (i) There exists a time instant \ud835\udf0f \u2208 R0+ such that the agents are classified into the players with Ki2 and the nonplayers from \ud835\udf0f",
            " Finally, let us consider the convergence speed. It has been shown in [4] that the coverage controller in (5) corresponds to the gradient descent algorithm for J. So, it follows from (i) and (ii) in the proof of Theorem 1 that the players converge as fast as the algorithm after a while. The detailed analysis will be performed as a future work. In this section, we verify the proposed controllers. In particular, numerical experiments with the standard images in [18] are performed. Consider the multi-agent system in Fig. 2, where n \u2236= 7500, S \u2236= [0, 400] \u00d7 [0, 400], F \u2236= [50, 350] \u00d7 [50, 350], and a \u2236= 10\u22124. The reference image r is Mandrill in Fig. 3(a). This is an eight-bit grayscale image, and thus r(p) \u2208 {0, 1\u2215255, 2\u2215255,\u2026 , 1}. The initial formation x0 is given randomly from the uniform probability distribution on (F\u29f5cent(F, 1))n. We use the proposed controllers in (2), (6), and (7) with k \u2236= 10 and c \u2236= 1.3 satisfying (11). Fig. 5 illustrates the time series of the resulting formations, where the solid squares represent the agents"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003117_s00542-020-04893-8-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003117_s00542-020-04893-8-Figure1-1.png",
        "caption": "Fig. 1 Two scotch-yoke mechanisms that have different yoke shapes",
        "texts": [
            " The Scotch Yoke mechanism was applied to the model to create the desired stiffness curve using simple linear springs. The described mechanism is commonly used for converting rotational motion into linear motion. Previous studies applied this mechanism to balancing applications or a series of elastic structures to take advantage of the ability to create gait cycles in robots (Shieh and Chou 2015; Ryder 2015; Ryder and Sup 2013). However, unlike other studies, changing the yoke shapes, as shown in Fig. 1, was the focus of this paper. When the disk is rotated by the motor, the roller fixed to a point on the disk also rotates in the direction of the disk rotation. At this time, the yoke is pushed by the roller as much as the roller moves in the horizontal direction, so that the linear motion part where the yoke is connected is translated horizontally. Here, if linear springs are installed on both sides of the yoke as shown in Fig. 2 and the motor shaft is connected to the input shaft of it, the design can have compliance characteristics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000900_s12206-014-0633-1-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000900_s12206-014-0633-1-Figure4-1.png",
        "caption": "Fig. 4. Mises stress distribution of plane (z = 0.25a): (a) \u03b8 = 37\u00b0; (b) \u03b8 = 60\u00b0.",
        "texts": [
            " Meanwhile, the stress changes greatly at or near the contact surface. Thus, the ranges of x, y axis are from -1.5a to 1.5a and the range of z axis is from 0 to 1.5a. Since the subsurface stress distribution is not symmetric, it is not suitable to display the subsurface in cylindrical coordinates. The streamlines of the von Mises stress at different subsurface plane is plotted in rectangular coordinate system. To illustrate the effect of contact angle, the von Mises stress distribution at subsurface x-y plane is shown in Figs. 4 and 5. From Fig. 4, it is found that the magnitude of von Mises stress is much higher when the contact angle increases. For example, the maximum von Mises stress is 2800 Mpa for \u03b1 = 60\u00b0, while it becomes 2200 Mpa for \u03b8 = 37\u00b0. Similarly, the shape of von Mises stress distribution is very like the shape of surface contact area. The shape of von Mises stress distribution tends to be elliptical as the contact angle decreases. Moreover, for increasing z depth, the value of stresses be- comes smaller, with the same proportional distribution, as described by Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001231_12.2040234-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001231_12.2040234-Figure1-1.png",
        "caption": "Figure 1. Electrowetting principle: (a) Initial condition of water drop at V=0; (b) Water drop shape after voltage applied; (c) Pin-holes in dielectric layer cause a direct contact between water and electrode layer, which leads to electrolysis that indicated by bubbles and forming",
        "texts": [
            "eywords: Electrowetting-on-dielectric, Parylene C, Al2O3, multi-layer insulator, sandwich-like multi-layer, breakdown voltage Electrowetting is a term referred to the ability of a liquid to change their wetting properties due to an applied voltage between the liquid and the surface as shown in Figure 1a and Figure 1b. Wide range of its possible application has made electrowetting technology gets a lot of researcher\u2019s attention. A number of researches have been reported in electrowetting application such as electrowetting display (e-paper)1, actuator2, photovoltaic3 and tunable liquid lens4. The basic of electrowetting technology was firstly introduced by Gabriel Lippmann5. The electrowetting behavior or the contact angle changes at a certain voltage is determined by Young-Lippmann equation \u03b3 \u03b8\u03b8 2 coscos 2CV Y += (1) Where \u03b8Y is the Young\u2019s or initial contact angle, C the capacitance per unit area, V the voltage applied to the system and \u03d2 is the interfacial tension between aqueous and ambient phase",
            " The insulator material will become electrically conductive if the applied voltage was exceeding that limit, resulting a dielectric failure or dielectric breakdown. Thinner layer has a lower voltage limit, enlarging the possibility of dielectric breakdown. In addition to the applied voltage limitation in electrowetting application, pinholes problem in a thin insulator layer also comes as a consideration. Even Parylene C, the most common moisture barrier, suffers for a severe pinholes problem when used as a thin layer6. Figure 1c visualizes how those pinholes serve as a pathway for the liquid through the insulating layer and go to the electrode layer, causing electrolysis and an early dielectric breakdown. To anticipate this problem, multi-layer insulator structure was employed to enhance the breakdown voltage6,7. When further increased in breakdown voltage is required, some simple approaches such as increasing the insulator thickness or adding more layers might be applied. However, those approaches will lead to an unwanted increase in operational voltage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003796_iros45743.2020.9341685-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003796_iros45743.2020.9341685-Figure2-1.png",
        "caption": "Fig. 2. Definitions of the angles.",
        "texts": [
            " The proposed motion planner estimates knee joint positions (as outputs) according to each estimated speed and gait percent. The speed and gait percent are estimated based on the thigh acceleration and angle (inputs) as observed in Fig. 1. To calculate the knee joint positions during the experiment, another IMU was attached to the shank of the subject to measure the shank angles \u03b8sh. This IMU is not required for the estimation task, but was required in order to obtain the corresponding knee joint positions \u03b8k during walking experiments (\u03b8k = 180\u2212(\u03b8sh\u2212\u03b8th), Fig. 2). These calculated knee joint positions, are then compared with the estimated knee joint angles (by the motion planner), to verify the estimation quality. To estimate knee joint positions, the proposed motion planner estimates the speed and the gait percent (Fig. 1). These steps are explained in detail in the next subsections. To estimate the walking speed, the concept of support vector machine (SVM) [20] has been used. SVM is a supervised machine learning approach which uses a set of training inputs for the learning process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001042_2014-01-1664-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001042_2014-01-1664-Figure2-1.png",
        "caption": "Figure 2. Side view of the engine showing inward and outward motion of the oil seals due to the eccentric motion of the rotor",
        "texts": [
            " IOC accounts for about half the total oil consumption, which is far greater than the necessary amount of oil to lubricate the contact between the oil seals and the side housing. A better understanding of the oil transport mechanisms through the oil seals would help reduce IOC, as the oil seals are the main barrier between crankcase oil and the combustion chamber. In particular, as the O-rings are sealing the oil seal grooves, oil must pass at the seal-housing interface during inward-outward motion of the seals created by the eccentric motion of the rotor, as shown in Figure 2. Details of the rotary engine kinematics are given by Yamamoto [1]. CITATION: Picard, M., Baelden, C., Tian, T., Nishino, T. et al., \"Oil Transport Cycle Model for Rotary Engine Oil Seals,\" SAE Int. J. Engines 7(3):2014, doi:10.4271/2014-01-1664. 1466 Apex seal dynamics and contact have been the focus of many studies, both from experiments (e.g. [2, 3]) and modeling (e.g. [4, 5]). However, oil seals have been the focus of only a few studies. Froede [6] shows the oil consumption for three oil seal configuration from the early development of the rotary engine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000408_1.4040999-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000408_1.4040999-Figure7-1.png",
        "caption": "Fig. 7. Schematic Experimental Set-Up for Mistuning Identification (see Beirow et al. [19])",
        "texts": [
            " Beirow et al. [19, 20] demonstrated an approach to determine experimentally blade individual eigenfrequencies. Therefore a blisk is excited by a miniature modal hammer while the vibration response is measured non-intrusively by a laser vibrometer. Further details are given in US Patent [21]. Summarizing, all blades except the excited one are detuned with additional masses to decouple one blade from the already disturbed cyclic symmetry. The miniature modal hammer excites sequentially all blades. Figure 7 according to Beirow [19] visualizes the schematic set-up. Note, that the illustrated blisk is not the investigated blisk in this paper. The effect of mass detuning is visible in comparison of Fig. 8 and Fig. 11. Without additional mass (Fig.8), there are at least two dominant resonant peaks and four barely visible conspicuous resonant frequencies related to six coupled modes of the blisk in the frequency range between 3500 Hz and 4500 Hz. With additional mass, there is only one resonant peak in this frequency range (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000834_2014-01-2064-Figure23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000834_2014-01-2064-Figure23-1.png",
        "caption": "Figure 23. The structure of clutch with small stiffness and large torsional angle",
        "texts": [
            " The torsional vibration is calculated by the collected pulse signals and gear numbers of the flywheel and transmission, then 2nd order torsional vibration signals at the flywheel and transmission are abstracted. The torsional vibration varying with the engine speed is measured at transmission input shaft for two clutch with stiffness values of 12.45 and 8N.m/degree respectively. For 4th gear and 5th gear, the resonance at transmission shaft happens at 1800 and 1700rpm respectively, as shown in Figure 22 and Figure 23. The subjective evaluation shows that there exists significant interior noise. The clutch with smaller stiffness reduces the torsional vibration magnitude significantly, which is consistent with the results from the clutch stiffness analysis and DOE analysis. According to the subjective driving results, torsional vibration and interior rattle noise are significantly reduced by only using the clutch with small stiffness and large rotation angle, especially at 4th gear and 5th gear. Figure 23 shows the clutch structure. The special structure increases the rotation angle and also reduces the stiffness. The following conclusions can be made according to the multi-body model analysis and test results. 1. The torsional modal frequency varies with the gear shifts for the drivetrain system. The modal frequency for 2nd and 3rd modes decreases with the gear shifts increase, while the 1st and 4th modal frequency increases with the gear shifts increase. 2. The forced vibration analysis shows that the resonant frequency of the transmission shafts is consistent with the system modal frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001316_066101-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001316_066101-Figure4-1.png",
        "caption": "Figure 4. Volume schematic diagram of V2.",
        "texts": [
            " We start with the right-angle hexahedral unit in order to obtain the porosity of the overall lattice structure. Let all cases of the right-angle hexahedral unit be constituted by AB, BC and BD, then disassemble the empty part of the hexahedral unit into the sum of several empty parts; refer to figure 3 for the schematic diagram of the disassembly. Disassemble the empty parts of a hexahedral unit into one small hexahedron, two top surface cuboids; two lateral surface cuboids; and two front surface cuboids. Figure 4 shows the sections of strut AB, strut BD and strut BC, where the sectional area of strut AB is SAB = h2h3, strut BD is SBD = h1h3, and strut BC is SBC = h1h2. There are N hexahedral units in the entire lattice structure, with O units along the X axis direction, P units along the Y axis direction, and Q units along the Z axis direction; thus N = O\u00d7 P\u00d7 Q. Vvirtual object = V1 + 2V2 + 2V3 + 2V4. (1) From the given conditions we obtain: V1 = (a\u2212 2h1)(b\u2212 2h2)(c\u2212 2h3). (2) As AB = a,BD = b, and using figure 4 we obtain: A\u0304B\u0304 = a\u2212 2h1 B\u0304D\u0304 = b\u2212 2h2 (3) and V2 = (a\u2212 2h1)(b\u2212 2h2)h3. (4) In the same way we obtain: V3 = (a\u2212 2h1)(c\u2212 2h3)h2, (5) V4 = (b\u2212 2h2)(c\u2212 2h3)h1. (6) As shown in figure 5, it is known that the lattice structure has N hexahedral units giving a volume NV1. In addition, as there are O hexahedral units in the X axis direction, P units in Laser Phys. 23 (2013) 066101 J Sun et al the Y axis direction and Q units in the Z axis direction, thus there are contributions of (Q + 1)OP V2; (O + 1)QP V3; and (P+ 1)OQ V4 and the total for all virtual objects is Vall virtual object = NV1 + (Q+ 1)OPV2 + (O+ 1)QPV3 + (P+ 1)OQV4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001803_978-3-319-07944-8_18-Figure18.9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001803_978-3-319-07944-8_18-Figure18.9-1.png",
        "caption": "Fig. 18.9 A three-dimensional representation of the resin bath (drum type)",
        "texts": [],
        "surrounding_texts": [
            "Using Improved Method Figure 18.12 shows the composite hollow shafts made by different types of raw materials (carbon, glass and kenaf fi bres) that were fabricated by fi lament winding method. It shows the carbon and glass composite shafts were fabricated using the 18 Filament Winding Process for Kenaf Fibre Reinforced Polymer Composites 380 S. Misri et al. 381 regular method of resin bath fi bre-dip type (as shown in Fig. 18.7 ), while the kenaf hollow shaft (Fig. 18.13 ) was fabricated using the improved method (resin bath drum type) (as shown in Fig. 18.8 )."
        ]
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure9-1.png",
        "caption": "Fig. 9. Hyperboloidal surface with embedded involutoid.",
        "texts": [
            " These constants are determined by recognizing that an involutoid is a curve embedded in a single sheet hyperboloid where the evolutoid is the throat as illustrated in Fig. 6 . The throat radius is the base radius of (6) and the generator angle is the angle of incidence \u03d1 of (7) . The normal N to the hyperboloidal surface (not the pitch surface) at is perpendicular to the generator through , thus m \u2261 \u03b1 \u03b2 = \u2212 N \u00b7 b N \u00b7 n = \u2212 ( \u2032 \u00d7 C \u00d7 ) \u00b7 b ( \u2032 \u00d7 C \u00d7 ) \u00b7 n (14) where the direction from E to is \u03b2( m n + b ) . Depicted in Fig. 9 is a hyperboloidal pitch surface along with a single sheet hyperboloid and an embedded involutoid. The intersection between the pitch surface and the hyperboloidal surface is a transverse curve. Also, the intersection between the pitch surface and an axial surface is a transverse curve. Each value of \u03b2 defines a point on the hyperboloidal surface. The perpendicular distance r from the central axis of the involutoid is r = \u221a ( + \u03b2m n + \u03b2b ) \u00b7 ( + \u03b2m n + \u03b2b ) \u2212 [ ( + \u03b2m n + \u03b2b ) \u00b7 C ] 2 . (15) Differentiating above w"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure3-1.png",
        "caption": "Fig. 3 Dual interlocking circular cutters",
        "texts": [
            " The two angles aa and a0 can be calculated based on the gear parameters aa \u00bc se 2ha tan an Re= cos dd a0 \u00bc tan 1\u00f0tan an= cos\u00f0aa=2\u00de\u00de 8< : (2) where an is the pressure angle of the gear, dd is the dedendum angle, Re is the outer cone distance, and se is the outer circular thickness. These parameters can be obtained using formulas listed in the AGMA standards [10]. Coniflex VR SBGs are manufactured by large-size plate cutters whose design details are described in Ref. [11]. In the dual interlocking cutters shown in Fig. 3, a plurality of blades is peripherally arranged on the cutter, thereby enabling higher productivity. To simultaneously cut both gear flanks, two identical cutters must be positioned in an interlocking arrangement. During the cutting process, each cutter group cuts only one slot at a time and then indexes to the next slot until all teeth are finished in what is termed a single indexing process. The cutters are inclined at a specific angle (e.g., 22 deg) to the generating plane, forming a tooth of the imaginary generating gear, whose cutters then infeed to the required tooth depth and generate the work gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002341_ijrapidm.2016.078746-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002341_ijrapidm.2016.078746-Figure16-1.png",
        "caption": "Figure 16 Bevel gear model with small surface blemishes",
        "texts": [
            " However, the graph shows that the angular tolerance is within 0.65\u00b0 for all the three angles. The bevel gear model possesses high level of rotational symmetry with smallest of differences in surface curvature. However, the model intricacy rests on more number of protrusions and holes. The prime objective was to examine whether the parts could be produced with adequate accuracy that would allow the gears to mesh. The developed models were free from any major issues; only small blemishes occurred on some teeth of the gears, as shown in Figure 16. Finally, the bevel gears having proper meshing of the teeth were assembled together. This section summarises various issues encountered during development of the proposed CAD parts using the ELAM R3D2 machine. Further, the possible causes and suitable corrective action to be taken for minimising these issues were proposed based on the experiments conducted and the observations made (Table 4). This table will provide a helpful lead for the users and machine developers to improve the machine further"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000528_978081000342.349-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000528_978081000342.349-Figure4-1.png",
        "caption": "Figure 4: Top: approximation of the rotor-bearing system, bottom: two",
        "texts": [
            " The gap function is given by \u210e( , ), is the radius of the bearing and the dynamic viscosity of the fluid. The shear stresses ( , ) are calculated by the assumption of NEWTON\u2019s hypothesis: = \u210e (\u03a9 \u2212 \u03a9 ) + \u210e2 , = \u210e \u03a9 \u2212 \u210e2 . resulting forces , , and , , in Cartesian coordinates and friction torques , can be calculated with , , = \u2212 , cos , / / , , , = \u2212 , sin , / / , , = , , ./ / A simple numerical model of the rotor system is used to calculate the rotor amplitudes and the rotational speed of the floating rings. Figure 4 shows the approximation of the rotor-bearing system. The rotor model contains four mass points on a massless, elastic shaft. The masses and are the masses of the real turbine wheel and the real compressor wheel. The masses , and , in the position of the journal bearings are chosen in a way that the centre of gravity of the rotor is maintained. The two floating rings are approximated by , and , . Each mass point has a translational degree of freedom, which is included in the complex state vectors = , , , , , , = , , , ",
            " The outer oil-whirl does not appear anymore within the investigated speed range after modifying the geometry of both journal bearings as mentioned above. The inner oil-whirl is followed by amplitudes at 0.75 \u210e , , which is closer to the experiment. The rotor enters the constant tone state at 0.7 \u2126 , and rotates with amplitudes at 1.6 \u210e , , which has a sufficient correlation to the experimental results. However, the numerical model shows a jump of the compressor\u2019s amplitude before leaving the constant tone in the range (0.55\u20130.65) \u2126 , . During this speed range, the rotor vibrates in the first bending mode shape, see Figure 4 bottom right. While in the speed range of the inner oil-whirl and the constant tone respectively, the rotor gets into a conical state of motion, see Figure 4 bottom left. The frequency 2 of the constant tone is better estimated with 0.49 to 0.52 \u2126 , . The rotational speed of the floating ring is still calculated too high. Comparison between the experimental run down and the numerically simulated run down is carried out using the speed ratio of the floating ring speed and the rotor speed. After determining experimentally the rotational speed of the floating ring using the high-speed camera, the ring speed ratio results in the relative speed range \u03a9 , /\u03a9 \u2248 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003248_j.cja.2020.06.030-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003248_j.cja.2020.06.030-Figure10-1.png",
        "caption": "Fig. 10 Parallel trimming of diamond wheel.",
        "texts": [
            " Through the axial and radical movement of the diamond dressing wheel and the rotational movement of the grinding wheel around its own axis to com- Please cite this article in press as: YANG S, CHEN W Modeling and experiment of g doi.org/10.1016/j.cja.2020.06.030 plete the trimming of the working face of the grinding wheel. The gear hob is then ground by the radial and rotational movement of the well-trimmed grinding wheel and the helical motion of the hob. This experiment uses the parallel trimming method of the diamond dressing wheel that reciprocates to trim grinding wheel in sequence of 1 to 7 as shown in Fig. 10. The tooth profile errorFfsof gear hob is the deviation of the actual position of a point on the cutting edge from the specified position on the envelope thread and the deviation value is defined in a measurement surface which is perpendicular to rinding wheel axial profiles based on gear hobs, Chin J Aeronaut (2020), https:// Table 2 Grinding wheel specifications and grinding parameters. Item Unit Value Diameter of grinding wheel mm 60 Abrasive material White corundum Abrasive diameter lm 300\u2013250 Grinding wheel rotation speed r/min 6000 Workpiece speed mm/min 4500 Grinding depth mm 0",
            " 13 which is obtained by previous algorithm that mainly take the rake face of the hob as the calculation basis to achieve the axial profile generation of the grinding wheel, we get the tooth profile errors for both flanks of hob A and hob B in Fig. 13 are 16.4 lm, 15.7 lm, 22.2 lm and 21.8 lm, respectively. It can be obviously seen that for the same hob, the tooth Please cite this article in press as: YANG S, CHEN W Modeling and experiment of g doi.org/10.1016/j.cja.2020.06.030 profile errors in Fig. 10 are much less than those in Fig. 13. As expected, the experiments of relief grinding demonstrate that in case of actual processing, the application of this method make the accuracy of hob tooth profile can better satisfy production requirements. ing experiment. rinding wheel axial profiles based on gear hobs, Chin J Aeronaut (2020), https:// 6. Conclusions The relief grinding process plays an extremely important role in the production of gear hob and the grinding accuracy will have a direct and profound influence on the hobbing accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003100_j.matpr.2020.04.476-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003100_j.matpr.2020.04.476-Figure5-1.png",
        "caption": "Figure 5. FEA Analysis for Gradient Porous Structure \u2013 Pore Volume Fraction 0.6-0.4",
        "texts": [],
        "surrounding_texts": [
            "The CAD model of the structures was generated using K3DSurf v0.6.2 software. Nine structures were generated in the software similar to ones shown in figure 2 (uniform porous structure) and figure 3 (gradient structure). Model size: 12.7 mm x 12.7 mm x 12.7 mm as per ASTM A370 Standard. 4. FEA Analysis A.S. Babu et al. / Materials Today: Proceedings 24 (2020) 1561\u20131569 1565 Table 2. Numerical Results of FEA analysis S. No. Porosity (%) Von- Mises Stress (MPa) Deformation (mm) Mesh 1. 10 103.25 0.02225 238298 2. 20 112.58 0.10079 248224 3. 30 142.14 0.10709 248893 4. 40 194.99 0.12469 228453 5. 50 205.25 0.13125 228453 6. 60 208.08 0.09437 224958 7. 70 214.74 0.10561 228467 8. 80 221.94 0.10716 224862 9. 90 230.18 0.13929 235128 Figure 4 and 5 shows FEA analysis for respective uniform and gradient structures. From table 2 where numerical results are tabulated, it is clear that Von-mises stress is inversely proportional to compressive strength. Structure with 90 % porosity is having maximum porosity as well as von-mises stress. Structure with 10 % porosity is having least von-mises stress. So to have a mixture of good strength as well as porosity of the structure, model with 60 % porosity is selected for fabrication."
        ]
    },
    {
        "image_filename": "designv11_30_0000542_analsci.31.591-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000542_analsci.31.591-Figure1-1.png",
        "caption": "Fig. 1 The coulometric microdevice for pesticide determination. (A) Exploded view of the device. (B) Top view of the device with the electrodes and the flow channel structure. (C) Picture of the device.",
        "texts": [
            " Commercial OP formulation (malathion) was purchased from Sumitomo Chemical Garden Products (Tsukuba, Japan). OP standard solutions were prepared with distilleddeionized water. For the analysis on the device, a solution containing ATCh and a solution containing OP and the enzyme were prepared. All reagents were of analytical grade. The microdevice consisted of a glass substrate with a threeelectrode system to measure charge generated accompanying the oxidation of TCh and a PDMS substrate with a flow channel structure (Fig. 1). The fabrication process was described in our previous work.8.10 The three-electrode system consisted of a platinum working electrode, an Ag/AgCl reference electrode, and a platinum auxiliary electrode. The working electrode consisted of a microelectrode array (40 thin strips of 600 \u03bcm \u00d7 400 \u03bcm) with a 130-\u03bcm inter-electrode distance (edge to edge). Two pinholes (30 \u03bcm in diameter) were formed on the silver layer of the reference electrode. AgCl was grown from the pinholes into the silver layer by applying a current of 50 nA for\u00a0 20 min in a 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000900_s12206-014-0633-1-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000900_s12206-014-0633-1-Figure6-1.png",
        "caption": "Fig. 6. Mises stress distribution of plane (y = 0): (a) \u03b8 = 37\u00b0; (b) \u03b8 = 60\u00b0.",
        "texts": [
            " 4, it is found that the magnitude of von Mises stress is much higher when the contact angle increases. For example, the maximum von Mises stress is 2800 Mpa for \u03b1 = 60\u00b0, while it becomes 2200 Mpa for \u03b8 = 37\u00b0. Similarly, the shape of von Mises stress distribution is very like the shape of surface contact area. The shape of von Mises stress distribution tends to be elliptical as the contact angle decreases. Moreover, for increasing z depth, the value of stresses be- comes smaller, with the same proportional distribution, as described by Fig. 5. Fig. 6 illustrates the von Mises stress in x-z plane when the contact angles are 37\u00b0 and 60\u00b0, respectively. From Fig. 6, we can conclude that the magnitude of von Mises stress is much higher when the contact angle increases. For increasing con- tact angle, it is found that the evolution of von Mises stresses meets well with the variation trend of maximum contact pressure, as described by Table 1. However, the shape of von Mises stress distribution is more like an ellipse when the contact angle decreases. Under the same normal load, the location of maximum von Mises stress is prone to be much deeper from the origin contact point when the contact angle increases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001689_jab.29.5.616-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001689_jab.29.5.616-Figure1-1.png",
        "caption": "Figure 1 \u2014 Incremental change in leg length ( dL  ) between two incremental time steps during stance depends on both the displacement of the COM ( COMds",
        "texts": [
            " F F \u00a0 0  =     (2) In this coordinate system, the radial leg vector L  can be described by its magnitude L and the enclosed angle \u03c6 between force the F  and the leg L  : L L cos sin  =       (3) Extending the work-loop concept to two-dimensional systems requires the estimation of the magnitude of the radial leg vector. Therefore, the total derivative of Eq. 3 is calculated to obtain the needed magnitude of the radial leg vector: dL L L dL L d = \u2202 \u2202 + \u2202 \u2202 (4) dL dL L d cos sin sin cos =       + \u2212      (5) The radial leg is defined as the vector pointing from the center of pressure (COP) to the COM. Hence, the change in leg length depends on the displacement of the COM, COMs  , and the displacement of the COP, COPs  (Figure 1): dL ds dsCOM COP = \u2212 (6) Combining Eq. 5 and Eq. 6 results in ds dL L d ds cos sin sin cos COM COP =       + \u2212      + (7) This equation describes the displacement of the COM with respect to the chosen coordinate system. Using Eq. 7, the work can be calculated as defined in Eq. 1: dW F dL FL d F ds\u00a0cos \u00a0 sinCOM COP = \u2212 + \u22c5 (8) Eq. 8 consists of three expressions: dW F dL\u00a0 \u00a0cos \u00a0Leg = (9) dW FL dsin = \u2212 (10) dW F dsCOP COP = \u22c5 (11)  ) and the displacement of the COP ( COPds  ). The vector of the ground reaction force is denoted with F  and the angle between the force vector F  and the radial leg vector L  is denoted with \u03c6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002870_s12206-020-0231-3-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002870_s12206-020-0231-3-Figure2-1.png",
        "caption": "Fig. 2. Vector diagram of a kinematic chain.",
        "texts": [
            " Each kinematic chain uses the RSS (R for revolute joint, S for spherical joint) construction, which has a motor, a reducer, a master arm, a slave arm, and two spherical joints. The master arms produce rotation movement driven by the motor. The slave arm is driven by the spherical joint mounted at the end of the active arm. The other end of the slave arm is connected with the moving platform. Through the six independent kinematic chains, the robot moving platform can achieve 6-DOF in space, namely, three translational motions and three rotational motions. Give that the six kinematic chains have nearly identical structures, the reference frame is defined as shown in Fig. 2. The coordinate frame O-xyz is attached to the static platform, and the coordinate frame P-uvw is attached to the moving platform. Therefore, the closed-loop position vector equation in the ith kinematic chain is expressed as 1 2 , = 1 ~ 6a l l i= - + +i bi i ir a TT u w r uur uur uur uur (1) where [ , , ] [cos( ),sin( ),0] * [cos( ),sin( ),0] [cos( ) cos( ),sin( ) cos( ), sin( )] T T i i T i i T i i i i i x y z b f f J J f q f q q = = = = = - i bi i i i r a T n n u r uur uur uur uur uur = \u03b11 \u03b12 \u03b13T T T T The inverse kinematics model can be obtained by Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002723_0142331219894867-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002723_0142331219894867-Figure3-1.png",
        "caption": "Figure 3. Magnetic Levitator. The obstacle rotates randomly and is placed on the path of the ball.",
        "texts": [
            " Hence, in order to achieve the widest region of attraction along with the smallest ultimate bound, we present the following optimization problem that also considers practical control issues explained in Theorem 2, Remark 1, and Remark 2 min : Pm i= 1 trace Lif g+v1d0 v2r s:t: : r \u00f8 d= 1,m \u00f8 1, d0. d= 1, 16\u00f0 \u00de 22\u00f0 \u00de, 24\u00f0 \u00de, 25\u00f0 \u00de, 35\u00f0 \u00de 38\u00f0 \u00de \u00f039\u00de where v1, v2 . 0 are real scalar as tradeoff coefficients. Illustrative example To illustrate the effectiveness of the proposed approach, we present a physically motivated nonlinear impulsive system. Consider the magnetic levitator system illustrated in Figure 3, which is mathematically described as (de Souza et al., 2014; Marquez, 2003) m\u20acy= km _y+mg lmli2 2 1+mly\u00f0 \u00de2 +f t\u00f0 \u00de, \u00f040\u00de where m= 0:068 Kg is the mass of the ball, g = g0 +Dg = 9:860:001 m=s2 is the gravity acceleration, km = km0 +Dkm = 0:001610 5 N s=m is the viscous friction coefficient, l=l0 +Dl= 0:4660:0001 H is the inductance, ml = 2 m 1 is the variation of the inductance, f t\u00f0 \u00de 2 0:001, 0:001\u00bd is the other kind of uncertainties (neglected forces), i is the electric current and y is the position of the ball"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000594_icarsc.2015.38-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000594_icarsc.2015.38-Figure1-1.png",
        "caption": "Fig. 1: The ALIV3 tilt-quadrotor prototype.",
        "texts": [
            " After the development [2] and construction [3] of the prototype ALIV, this work now seeks to guarantee a trustworthy platform and progress towards the tilt-quadrotor control. With that purpose in mind, and condensing the main information required to build a simulator, this paper presents the tilt-quadrotor concept and corresponding dynamic model, including the actuators and sensors models. The procedures and results of the experimental identification of the model parameters of the ALIV3 prototype (see fig. 1) are also described. To conclude, final remarks and future guidelines are drawn. This section introduces the tilt-quadrotor ALIV3 prototype and its concept. The ALIV3 dynamic model, including the model of the existing sensors and actuators, is also presented. The tilt-quadrotor ALIV3 platform (see fig. 1) consists of a structure with a central core and four arms, each with a motorpropeller set. Two opposing arms have fixed motor-propeller sets, while the other two arms have swivelled motor-propeller sets. The tilting motions are the result of the action of four servos. The two servos located in the swivel arms control the roll-tilt (\u03c6ti ) - see fig. 2a. The two other servos, located in the central core, are responsible for the pitch-tilt (\u03b8ti ) of the rotors (fig. 2b). Besides the pitch-tilt servos, the central core also allocates the battery, an ArduPilot APM1 (Arduino-based processor for flight control, including 3-axis accelerometer, gyroscope and 978-1-4673-6991-6/15 $31"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000849_powereng.2013.6635624-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000849_powereng.2013.6635624-Figure1-1.png",
        "caption": "Fig. 1. Two-dimensional FE Model",
        "texts": [
            " They study about starting synchronous machines with frequency converters and make a statement about the thermal load of damper bars. Their analysis is based on the two reaction model by R.H. Park, which is based on certain assumptions and simplifications. With the help of up to date finite elements and time stepping analysis it is possible to get a closer look on this topic. The electro-mechanical model is built up with the software ANSYS Mechanical APDL. It is a two dimensional crosssection of the generator. Figure 1 shows the whole FE-Model. By using symmetrical properties the model can be reduced to one-fourth of its earlier size. Effects at both ends are negligible. The gap between rotor and stator is separated into two areas. This is necessary to simulate a mechanical rotation. Every step in the transient simulation results in a new coupling with constraint equations between the nodes at the center lines. The rotating angle for every step can be calculated by using the electromagnetic torque and a simple torque equation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000648_tmag.2013.2290101-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000648_tmag.2013.2290101-Figure9-1.png",
        "caption": "Fig. 9. Finite element mesh in the cross section of the induction motor. (a) Rotor and stator. (b) Rotor (zoomed-in-view). (c) Stator (zoomed-in-view).",
        "texts": [
            " (17) Let us denote the air-gap width by \u03b4 = Rs \u2212 Rr . For typical electrical machines, \u03b4/Rr 1. For instance, for the test case presented in Section VIII ratio \u03b4/Rr \u2248 6 \u00b7 10\u22123. In this case, we may approximate |c| by \u03d1Rr/\u03b4. The maximum condition number is obtained for \u03d1 = /2. Note that for increasing \u03d1 , we have cond\u03bc x\u0302 = O ( \u03d1Rr \u03b4 )4 . (18) In this section, the start-up of an induction motor under no-load conditions is examined. A three-phase cage induction motor is considered, whose cross section can be observed in Fig. 9. Some parameters of the test motor are collected in Table I. The test case is a slightly simplified version of the 15 kW motor presented in [19, Sec. 3.3.4] and in [20], respectively. Note that, we aim at discussing the metricbased treatment of motion rather than producing results to be compared against the data in the references. The following modeling assumptions regarding the details of the machine have been made. The magnetic field in the motor is assumed to be 2-D, end effects are neglected, as well as the skew of the slots"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure12-1.png",
        "caption": "Fig. 12 Total equivalent elastic strain in socket and hair",
        "texts": [
            "42; hair density\u00bc 1100 kg/m3 and Poisson\u2019s ratio\u00bc 0.38. In this study, the deformation of the hair and socket under incremental line loads on the hair is studied. Under such load, the hair is expected to be deflected from the loading direction until it hits the socket edge. As the hair deflects along the loading direction, the socket structure will also deform. However, the hair and socket deformations are expected to be different during the precontact and postcontact periods between the iris and the hair. Figure 12 shows a typical contour plot of total strain in the socket and hair under the applied line load on hair. It is clear from the figure that the equivalent elastic strain is the highest in the belt and skirt. The von-Mises stress is also observed to be the largest Journal of Biomechanical Engineering AUGUST 2016, Vol. 138 / 081006-5 Downloaded From: http://biomechanical.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jbendy/935362/ on 03/13/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use in the same regions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001303_1.4024087-Figure20-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001303_1.4024087-Figure20-1.png",
        "caption": "Fig. 20 Standard single land journal bearing",
        "texts": [
            " To assess the importance of the moment stiffness terms for journal bearings, we evaluate representative stiffnesses using a numerical analysis. A journal bearing with a length-to-diameter ratio of about 0.62 (similar to those in the NASA GRC test gearbox) was modeled and rotated slightly about the shaft in both transverse directions. The updated pressure distributions were computed for each rota- tion angle and integrated to compute the resulting forces and moments. The moment stiffnesses were computed using a simple finite difference calculation, where Kmoment ffi Dmoment Dangle Figure 20 shows a standard journal bearing with the majority of the film impedance concentrated in an annular section reacting against the static load and Table 5 shows the properties for this demonstration analysis. The fluid film pressure distribution for the journal bearing is shown in Figure 21, revealing the peaked behavior of the pressure field. When the shaft is rotated about the x(a) and y(b) axes, as demonstrated in Figure 22, the film pressure distributions also shift slightly, leading to a moment stiffness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002390_gt2016-57410-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002390_gt2016-57410-Figure4-1.png",
        "caption": "Figure 4. Rear bearing configuration",
        "texts": [
            " In this way, the seals before the impulse wheel are characterized by high cross-coupling stiffness coefficients. 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89517/ on 03/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use For the initial configuration, the turbine rotor, shown in Figure 2, was supported by two tilting pad journal bearings and a double tilting pad thrust bearing; 3D models and cross sections of the bearings are shown in Figure 3 and Figure 4. The shaft sizes and geometric characteristics of the journal bearings are listed in Table 1. The front bearing, located at the steam admission side of the machine, was a horizontally split journal bearing combined with a selfequalized double thrust bearing with direct lubrication, arranged in a back to back configuration between the two thrust collars. The tilting pad thrust bearings consisted of 6 offset pivot steel backed pads on each thrust face, supplied with oil using bi-directional spray nozzles between the thrust pads"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003588_lra.2020.3036569-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003588_lra.2020.3036569-Figure1-1.png",
        "caption": "Fig. 1. Experimental setup with (a) highly-flexible object and 4 robots to visualize impact of transport speed on deformation and (b) local force sensing mechanism on the grasped object.",
        "texts": [
            " For example, when transporting flexible objects, such as large uncured-composite aircraft wings in a manufacturing setting, large deformations can lead to damage. Therefore, this letter selects the parameters of the accelerated gradient-based update approach to increase the transport speed without increasing the deformations of the flexible object. An advantage of the proposed gradient-based approach is that it can be implemented with only local force information using an accelerated delayed self reinforcement (A-DSR), without the need for additional information. An example of a robotic team, shown in Fig. 1, is used to comparatively evaluate the performance, with and without the proposed A-DSR approach. The transported object is chosen to be highly flexible to help 2377-3766 \u00a9 2020 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: Raytheon Technologies Corporation. Downloaded on December 19,2020 at 10:19:23 UTC from IEEE Xplore. Restrictions apply",
            " (11) with the update parameters (\u03b3, \u03b21, \u03b22) selected to minimize the transport time with the same deformation as in Case 1. In this experiment, a transport task along a single axis, i.e., zk = [0, 0, zk] T , fk = [0, 0, fk] T with zk \u2208 R, fk \u2208 R, is considered to demonstrate the efficacy of the proposed approach. The experimental system consists of four custom mobile robots shown in Fig. 4. Each robot is equipped with magnetic encoders on the wheels for absolute position feedback, a micro-controller for on-board computation of control inputs and load cells as a force sensing mechanism, as shown in Fig. 1. Each robot has a grasping hook on the front that enables attachment to the flexible object. To visualize the deformation during transport, the object is chosen to be highly flexible, comprising of a coiled spring with diameter of 1.30 cm and length of 90 cm. The schematic robot transport is shown in Fig. 4. Each robot is connected to its nearest neighbor, and only the leader (robot 1) has knowledge of the desired position illustrated by the virtual robot shown in pink. The effective stiffness coefficients, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000900_s12206-014-0633-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000900_s12206-014-0633-1-Figure1-1.png",
        "caption": "Fig. 1. Stranded-wire helical spring.",
        "texts": [
            " This paper attempts to calculate the contact area and pressure of the elliptical contact between adjacent wires during working process based on the Boussinesq potential functions and elastic half-space model. In addition, these surface contact quantities are used to expand the calculation of subsurface stress relevant to elliptical contact. Further, the present work provides support for the investigation with respect to the mechanism on fatigue, wear of the stranded-wire helical spring. In recent years, Jie Zhou et al. [5] formulated a mathemati- cal model for stranded-wire helical spring, as shown in Fig. 1. The normal contact force between two adjacent wires is obtained via theoretical calculation when the mass impacted spring with a certain velocity [9]. It is well known that the contact angle between two steel wires is less than 90\u00b0, which brings an elliptical contact during working process. In this section, an analytical method is developed to calculate the contact angle between two adjacent wires. In rectangular coordinate system, the three axis coordinates of the central line of a selected wire are as follows: ( cos )cos sin sin sin ( cos )sin sin sin cos tan cos sin x R r r y R r r z R r j b a j b j b a j b b a a j = + + = + - = + (1) where R is center diameter of stranded-wire spring, r is center distribution radius of strands, \u03b1-helix angle of stranded-wire spring, \u03b2- the twist angle of the strand, \u03c6- rotating angle of centre line of a wire around centre line of the strand"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001438_cdc.2013.6760540-FigureI-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001438_cdc.2013.6760540-FigureI-1.png",
        "caption": "Figure I. Quadrotor reference frame and propeller configuration.",
        "texts": [
            " In addition, this step allows to establish the range of operation for the internal variables \u00a2 and (). This range is important since the QL approximation was derived according \u00a2 \"'\" 0 and () \"'\" 0 . Thus, it is relevant to elucidate the extension of the range where the QL model is still accurate. This range is shown in Fig. 4. In order to further elucidate the differences between the variables, the error between the models was calculated and presented in Fig. 5. In all cases it is assumed that the TNL model is the most accurate. From this fIgure it is clear that the level of error for the MNL is due to the numerical resolution of the numerical solver. On the other hand, the error of the QL approximation is due to the linear approximation, nonetheless it is very low. The validation for this operating range is completed by comparing the inputs of the models. That is, the control effort of the controller. This comparison is required in order to reduce the effect of the closed loop controller over the comparison. The idea behind this is: \"if the outputs of two systems are the same and the inputs are the same, then the systems are equivalent within the operating range\""
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure15.9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure15.9-1.png",
        "caption": "FIGURE 15.9 (a) Cross section between r and z and (b) perspective view of concentric cylinder structures.",
        "texts": [
            " The structure is infinitely long in Z, and we have been dealing only with a cross section of it. Since such a structure with no Z dependence (i.e., uniform) can be described entirely by voltages and fields that are only functions of X and Y, a 2d analysis finds everything there is to know about the structure. 15.3 Axisymmetric Structures 337 Voltages, electric field components and energy density are functions of X and/or Y only and are valid for any value of Z. Total energy and capacitance values are calculated as joules and farads per unit length. Now consider Figure 15.9. 338 Some FEM Topics Figure 15.9a is identical to Figure 15.7b except that the axes are now labeled (r,z) rather than (x,y). The structures represented, however, are not the same. Figure 15.9b shows the axisymmetric structure described by Figure 15.9a. There are two important points to make here: 1 Figure 15.9 shows the entire structure being modeled, not just a piece of it. It is a 3d structure shown in its entirety with both cylinders shorted at both ends. The capacitance between the inner and outer cylinders and the total stored energy (for some prescribed electrode voltages) are measured in farads and joules, respectively. 2 The location of z = 0 for the structure is arbitrary, as is the location of x = 0 and y = 0 in rectangular coordinates insofar as the properties of the structure are concerned. In other words, Figure 15.9b may be moved (caused to slide) up or down the z axis without changing any structure properties. On the other hand, the r axis is fixed. In order to retain the axisymmetric status, there can be no \u03c6 dependence, and the structure in Figure 15.9b must always be fully describable by its cross section (Figure 15.9a). Deriving the formulas for the coefficient matrix for an axisymmetric structure using triangles with the shape functions defined in Chapter 13 duplicates the analysis of Chapter 13, except that X is replaced with r and Y is replaced with z, up to the energy expression [equation (13.10)]: U = \u03b5 2 \u00f0 \u00f0 T E2dx dy =E2AT \u00f015:2\u00de U = 2\u03c0\u00f0 \u00de\u03b5 2 E2 \u00f0 \u00f0 T r dr dz \u00f015:3\u00de Evaluating the integral in equation (15.3) is not a trivial task. Fortunately, it is an already accomplished and well-documented task. If we put aside the physics behind equation (15"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000943_2669047.2669054-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000943_2669047.2669054-Figure3-1.png",
        "caption": "Figure 3: The internal structure of the retractable stylus.",
        "texts": [
            " Force feedback from display consist of a force in a direction of a shrinkage of the stylus and a friction force in parallel to the display surface. The friction force is the opposite direction to the direction of movement of the stylus. The friction force changes according to the frequency of vibration. Changing the frequency of given vibration enable user to get some force feedback. Shrinkage of the stylus enables a user to feel macroscopic force from virtual objects, and oscillation of the voice coil motor enables a user to feel microscopic force from surface\u2019s texture of virtual objects. Figure 3 shows an internal structure of the stylus proposed in this paper. A DC motor set in the stylus is employed to shrink and extend the stylus. A gear is attached to a shaft of the DC motor. A pressure sensor is attached to a tip of the stylus to measure the pressing force given by the user. And a voice coil motor is attached to the stylus to display the user the tactile sensation of the virtual object\u2019s surface by oscillating. Figure 4 shows the flow of processing to change the length of the stylus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000896_s00158-014-1215-7-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000896_s00158-014-1215-7-Figure2-1.png",
        "caption": "Fig. 2 a Particle movement in a basic PSO, b Particle movement with digital pheromones",
        "texts": [
            " However, using just two components potentially impedes desirable exploratory characteristics of the swarm in n-dimensional design spaces. Also, poor locations specified by pBest and gBest in the initial stages of optimization offsets the swarm from attaining the neighborhood of the optimum solution. Previous work by the authors addressed these two issues through the implementation of digital pheromones as an additional velocity vector component within PSO for improved solution accuracy and efficiency (Kalivarapu et al. 2008). Figure 2a displays a scenario of a swarm member\u2019s movement whose direction is guided by pBest and gBest alone. Pi-1 denotes the position of a particle in the previous iteration and hence shown in gray. With a velocity of Vi-1, the particle\u2019s position is updated to Pi. If c1 >>c2, the particle is attracted primarily towards its personal best position. On the other hand, if c2 >>c1, the particle is strongly attracted to the gBest position. In this scenario as presented in Fig. 2a, neither pBest nor gBest leads the swarm member to the global optimum, at least, not in this iteration adding additional computations to find the optimum. Figure 2b shows the effect of implementing digital pheromones into the velocity vector. An additional target pheromone component causes the swarm member to result in a direction different from the combined influence of pBest and gBest thereby increasing the swarm diversity and hence the probability of finding the global optimum. Pheromones are generated by the entire swarm, so using them incorporates previously unused information into the algorithm. The method initialization is similar to a basic PSO except that 50 % percent of the swarm within the design space is randomly selected to release pheromones in the first iteration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003771_ssci47803.2020.9308433-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003771_ssci47803.2020.9308433-Figure2-1.png",
        "caption": "Fig. 2: Illustration of a \u201cspider\u201d robot that contains one Core module, eight Brick modules, and eight Joints.",
        "texts": [],
        "surrounding_texts": [
            "Our robot system is based on RoboGen; the bodies of the robots are modular, composed of three types of modules including a joint module actuated by a servo motor [1]. The robot\u2019s brains (controllers) have a network structure where each joint of the body has a coupled oscillator to drive it and where neighboring joints are connected."
        ]
    },
    {
        "image_filename": "designv11_30_0002737_s11182-020-01877-z-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002737_s11182-020-01877-z-Figure1-1.png",
        "caption": "Fig. 1. Schemes of electron beam additive manufacturing (a) and of the FSP (b) comprising electron beam gun 1, wire feeder 2, wire 3, printed product 4, substrate 5, table 6, electron beam 7, shoulder of the FSW tool 8, probe of the FSW tool, stir zone 10 and 11, and workpiece 12. Here arrows I and II indicate the processing directions.",
        "texts": [
            " After that the plates were subjected to friction stir processing on an experimental stand to adjust the FSW modes. The FSW tool applied for welding of plates 1.5 mm thick was used for the FSP; therefore, a pass over the base metal simulated obtaining of a welded joint with thickness of 1.5 mm by the FSW method, except for the presence of the joint before welding. The FSP modes are presented in Table 2. The schemes of processes of obtaining workpieces and FSP plates for sample production are illustrated by Fig. 1. Processing was performed in two directions \u2013 along (I) and across (II) the material layer deposition direction (Fig. 1\u0430). After processing, the templates for metallographic studies were cut from the processed samples in the cross section of the stir zone and the tensile test coupons were cut along the stir zone (Fig. 2). The mechanical properties of the stir zone formed by the FSP were investigated. The coupons for tensile tests represented dog bone shaped specimens 1.5 mm thick cut by the electrical dischage machine from the plate side subjected to FSP"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001548_iciea.2015.7334299-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001548_iciea.2015.7334299-Figure6-1.png",
        "caption": "Fig. 6. Two-dimensional model of motor",
        "texts": [
            "[10], the whole eddy current loss on one PM can be obtained (11) Where is the axial length of PM, is the radial width of PM, is the radial thickness of PM. Equation(11) shows eddy current loss is proportional to the square of alternating frequency and the square of maximum flux density , is inverse proportional to the resistivity , and related to the structural parameters of PM. B. Simulation model A Two-dimensional model of FSCW-PMSM with 24-slot 16-pole is built by using FEM. On the basis of cycle symmetry, the whole model is divided into 1/8, and it is shown in Fig.6. Considering different area of the motor and the influence of skin effect, the grid subdivision shown in Fig.7. Fig.8 and Fig.9 represent the distribution of the flux density and the magnetic field lines on the motor model. The PM eddy current loss under rated-load condition is shown in Fig.10. It fluctuates periodically, because it changes with the change of the stator coil position. 1248 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) C. Eddy current loss of main harmonics The fundamental harmonic of air gap field has the same speed as rotor, so it does not produce eddy current on PMs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001460_j.ijsolstr.2013.05.004-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001460_j.ijsolstr.2013.05.004-Figure3-1.png",
        "caption": "Fig. 3. Typical traction distribution s(q) obtained after a series of slip reversals and correspondent memory diagram. Points b1\u2013b3 do not move, point b shifts to the left as slip propagates towards the center.",
        "texts": [
            " a situation when the rate of the contact radius a increase is larger than the rate of slip inward propagation, so that slip does not develop at all. Thus, in our case, the solution always remains hysteretic since slip is always present in the outer annulus b < q < a. A chain bm (m = 1..M) of previous stick\u2013slip boundaries determines the memory effects in the system. The memory diagram in that case is defined as a sequence of points bm and the point b, supplemented by alternating signs Im (Im = Im 1) and the current sign I (see Fig. 3). The solution to the problem depends on whether some memory points are present in the diagram (M > 0) or not (M = 0). In the latter case corresponding to the initial hysteresis curve, the result has already been obtained in Section 3; here we rewrite it as d \u00bc d\u00f0b\u00de () b \u00bc d 1\u00f0d\u00de T \u00bc T\u00f0b\u00de ( ; \u00f027\u00de where d\u00f0b\u00de and T\u00f0b\u00de are given by Eqs. (10), (18): d\u00f0b\u00de \u00bc l \u00f02 m\u00de\u00f01\u00fe m\u00de E Z a b qr\u00f0a;q\u00dedqffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi q2 b2 q ; \u00f028\u00de T\u00f0b\u00de \u00bc 4Eb \u00f02 m\u00de\u00f01\u00fe m\u00de d\u00f0b\u00de 4l Z a b qr\u00f0a;q\u00de p 2 asin b q dq: \u00f029\u00de The inverse function d 1\u00f0d\u00de in Eq",
            " Solution (27) is written for increasing and positive d and can be easily modified to include decreasing and negative d as well, b \u00bc d 1\u00f0jdj\u00de d 1\u00f0Id\u00de T \u00bc I T\u00f0jbj\u00de I T\u00f0Ib\u00de ( : \u00f030\u00de Note that for the initial hysteresis curve increasing argument means positive, and decreasing argument means negative. Leaving the initial hysteresis curve leaves also some residual traction in the contact zone. It can be taken into account by adding a factor of two into the solution, d \u00bc dM \u00fe 2I d\u00f0Ib\u00de () b \u00bc d 1\u00f012 I\u00f0d dM\u00de\u00de T \u00bc TM \u00fe 2I T\u00f0Ib\u00de ( ; \u00f031\u00de where the extremum point (dM, TM) is known. The fact that slip inward propagation not only \u2018\u2018writes\u2019\u2019 the traction s\u00f0q\u00de \u00bc ljr\u00f0a;q\u00dej but erases the residual stress is reflected in the memory diagram (Fig. 3) as grey filling. In a more general case then overloading is possible, one has to introduce more filling styles in order to reproduce various regimes of the system\u2019s evolution. Thus we have found the solutions corresponding to a given memory diagram. These expressions Eqs. (30), (31) have to be supplemented by an algorithm controlling the evolution of memory diagrams. During the evolution two events may happen: (i) passing an extremum results in the creation of a new memory point, and (ii) slip inward propagation i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000482_0954406215590170-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000482_0954406215590170-Figure3-1.png",
        "caption": "Figure 3. Dynamic milling cutting model.19",
        "texts": [
            " All these matrices are given in Chaari et al.9 {Fext(t)} represents the vector of forces applied to the system. It is expressed by: Fext t\u00f0 \u00de \u00bc 0, 0,Tm, 0, 0, 0, TL, 0f gT \u00f05\u00de where Tb and sb are torque at break-down (at maximum torque) and the slip, a and b are constants characterizing the asynchronous motor and s is the proportional drop in speed given by: s \u00bc \u00f0Ns N1\u00de Ns \u00f07\u00de where Ns is the synchronous speed. To model milling cutter with N teeth (no helix angle), two degrees of freedom model is adopted19 (Figure 3). Forces induced by cutting process cause a dynamic response characterized by displacements x in the feed direction (X) and y in the feed direction (Y). The coordinate transformation vi\u00bc x sin( i(t)) y cos( i(t)) allows carrying x and y to the ith rotating tooth in the radial or chip thickness direction. Here, i(t) is the instantaneous angular engagement of tooth (i) measured clockwise from the normal axis (Y). This angle is time varying since the spindle rotates at the angular speed of so that i(t)\u00bc t"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002923_0954406220914341-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002923_0954406220914341-Figure3-1.png",
        "caption": "Figure 3. (a) Schematic view of an LRG joint; (b) A-A cross-section of ball\u2013raceway joint surface.",
        "texts": [
            " The six load types applied to the LRG include the tangential load F1 oriented along the X-axis, the normal compressive load F2 oriented opposite to the Y-axis, the normal tensile load F2 oriented along the positive Y-axis, the pitch moment F4 rotated around the X-axis, the yaw moment F5 rotated around the Y-axis, and the roll moment F6 rotated around the Z-axis. The LRG pair is composed of a slider and a rail, and the joint between one slider and rail is known as the LRG joint cell. An LRG joint cell contains numerous ball\u2013raceway joint surfaces. The ball\u2013raceway joint surface can be classified as the joint surface between the ball and the slider raceway and is referred to as the slider ball\u2013raceway joint surface; the joint surface between the ball and rail raceway is known as the rail ball\u2013raceway joint surface and is illustrated in Figure 3(a). The slider ball\u2013raceway joint surface is taken as an example for the stiffness analysis of the LRG joint in this section. Under the external loads, the force balancing equation of the LRG joint can be written as Ff g \u00fe OFR \u00bc 0 \u00f09\u00de where Ff g \u00bc F1, F2, F3, F4, F5, F6f gT, F3 \u00bc 0, and OFR is the reactive force matrix of the LRG joint under the external loads. The reactive force matrix of the LRG joint can be determined as follows OakiJTo \u00bc OROaki 0 O Oaki OROaki OROaki \" # \u00f011\u00de O Oaki \u00bc 0 O Oaki3 O Oaki2 O Oaki3 0 O Oaki1 O Oaki2 O Oaki1 0 2 64 3 75 \u00f012\u00de where k is the column number of the rolling ball, (k\u00bc 1 to 4), I is the total number of rolling balls per column, OakiJTo is the transformation matrix, and OakiFakiR is the reactive force matrix of the ith slider ball\u2013raceway joint surface of the kth column"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000648_tmag.2013.2290101-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000648_tmag.2013.2290101-Figure8-1.png",
        "caption": "Fig. 8. Cross section of a typical electrical machine, the air gap is exaggerated. Rotor domain r , stator domain s , air-gap domain a . We assume that the air gap domain is split into two subdomains by a circular cylinder l (lock step), coaxial to the rotation axis of the machine. The finite element mesh on the cylinder shall be invariant with respect to rotation by lock-step angle .",
        "texts": [
            " The analytical solution, independent of z\u0302 = x\u03023, is \u03c6 = g(r\u0302)\u03c6\u0302 + ( 1 \u2212 g(r\u0302) ) (\u03c6\u0302 + \u03d1) (14) where the expression for g(r\u0302) is g(r\u0302) = ln r\u0302 Rr ln Rs Rr . The resulting transition map is shown in Fig. 7. In this case det J = 1 holds everywhere for all \u03d1 . Let us now examine in more detail tangible engineering examples. We aim at testing our ideas against the so-called lock-step or sliding surface method [15], [16]. Consider an electrical machine cross section and the corresponding finite element mesh according to Fig. 8. With the standard lock-step method, rotation of the rotor is divided into discrete steps of size . However, since the finite element mesh is constructed in the x\u0302 system, rotation does not imply that the mesh needs to be changed at every time step. Accordingly, the lock-step method is easily relaxed to allow any time steps, and this enables one to employ adaptive time-integration techniques. The theory itself admits mechanical rotation angles up to \u00b1\u03c0 , but for numerical reasons, standard finite elements limit this in practice",
            " This motivates a combined approach, which we will call by name chart-metric method. Section VII-A explains how adjustable steps can be introduced in the lock-step method. Section VII-B elaborates on large rotations and the related numerical issues. A criterion for the choice of will be discussed and presented in Section VIII-B. A description of how the presented approach could be implemented in a finite element software system is given in Appendix. A. Introduction of Adjustable Steps into Lock-Step Method The standard lock-step method requires a situation similar to Fig. 8. The finite element mesh shall be prepared in such a way that the circular cylinder l (approximately) consists of a union of element interfaces. Moreover, the mesh on the cylinder shall be invariant with respect to rotation around the machine axis by lock-step angle . In other words, mesh reconnection at l with angular displacements that are integer multiples of is supported. This setting serves as our starting point. With the chart-metric method, small angular displacements |\u03d1| < are treated by changing the metric of the coordinate system, while angular displacements greater than trigger a mesh reconnection at l ",
            " The test example demonstrates some practical advantages of such an approach when solving numerically boundary value problems modeling rotating machines. APPENDIX FINITE ELEMENT IMPLEMENTATION OF CHART-METRIC METHOD We describe how the chart-metric method presented in Section VII-A could be implemented in a finite element software system. We present a high abstraction procedural description of the most important steps, with some existing time-stepping scheme in mind. Derivation of more specific (pseudo)code should be straightforward. Before execution of time-stepping, perform the following tasks. 1) Prepare a mesh according to Fig. 8. Assign N = 2\u03c0/ and k = 0. Index k is used to keep track of the mesh connection. 2) Prepare data structures to solve the boundary value problem u = 0 with Dirichlet conditions on the boundaries s , r . Note that the boundaries s , r need not to be circular cylinders. The remaining boundaries of the domain a receive homogeneous Neumann conditions. Given the desired rotation angle \u03d1 , perform the following steps. 1) Determine index j = argmin j\u2208[0,N\u22121]|(\u03d1 \u2212 j ) mod 2\u03c0 |. If j = k, jump to step 3). 2) Mesh reconnection: reconnect the meshes with rotor rotated by j , relative to discretized position. Let k = j . 3) Assign \u03c8 = \u03d1 \u2212 k (21) the reduced rotation angle. Note that |\u03c8| \u2264 /2. 4) Coordinate generation: solve the boundary value problem u = 0, in a u = 0, on s u = \u03c8, on r \u23ab \u23aa\u23ac \u23aa\u23ad numerically; compare to Fig. 8. This yields the nodal values of u in \u0304a . Compute the gradients (\u2202u/\u2202 x\u0302, \u2202u/\u2202 y\u0302, \u2202u/\u2202 z\u0302). The boundary value problem is obtained from the Laplace problem for \u03c6 according to Section VI by the ansatz \u03c6 = \u03c6\u0302 + u. 5) Jacobian matrix computation: form the matrices R, R\u2032, S, J at the integration points of the finite elements, according to (16) and (17), respectively. 6) Metric modification: depending on the formulation, replace the air-gap permeability \u03bc0 by the matrix \u03bc x\u0302 according to (12), or reluctivity 1/\u03bc0 by matrix \u03bd x\u0302 according to (13), respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002540_ssrr.2016.7784286-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002540_ssrr.2016.7784286-Figure8-1.png",
        "caption": "Fig. 8. Distance to move in the normal climbing phase",
        "texts": [
            "00 \u00a92016 IEEE 115 where \u03b80p is the pitch angle of the robot body when the front subtrack makes contact with the upper floor after the robot falls down forward. Thus, it satisfies( lm 2 \u2212 t0 ) sin \u03b80p + r cos \u03b80p \u2212 lf sin(\u03b8f \u2212 \u03b80p)\u2212 r = 0. (11) As we assume that the robot has no external sensors in this research, the phase transition from the normal climbing phase to the pitch down phase is conducted when its running distance D in the normal climbing phase exceeds a fixed value, as shown in the following Equation: D \u2265 d+ (n\u2212 1)p, (12) where n is the number of steps. d is a dimension indicated in Fig. 8, which according to the geometric condition is determined by d = (r+yc) tan \u03b8s+ p tan \u03b8s \u2212 r(1\u2212 cos \u03b8s) sin \u03b8s \u2212 lm 2 \u2212xc. (13) V. VERIFICATION TEST We verified our proposed method described in the previous section by conducting tests with a tracked robot on mock-up stairs. In this section, the procedures, results, and discussion of these verification tests are described. 1) Tracked robot: We used the tracked robot Kenaf [6][7] in our laboratory for verification tests (Fig. 9). The Kenaf is a 6-DOF tracked robot: two main tracks for traversal, and four subtracks, which are located on both sides at the front and rear of the robot, and can be controlled independently"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000763_icengtechnol.2014.7016810-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000763_icengtechnol.2014.7016810-Figure3-1.png",
        "caption": "Figure 3. Configuration, inertial and body fixed frame of the quadrotorl61",
        "texts": [
            " The Nonlinear feedback controller proves the ability to stabilize both roll and pitch angles at the same time while classical PD controller, with low pass filter, can only deal with one of these two angles at a time, as shown in the result and analysis section. II. MATHEMATICAL MODEL OF QUADROTOR The quadrotor system mathematical model was the focus of many researchers in the last decade such as [1], [2], [4] and [5].Inertial frame outside the model at a fixed point and body frame with its origin located in the quadrotor center of mass were assumed to model the quadrotor system, as shown in Fig. 3 Quadrotor orientation is defmed by three Euler angles namely roll angle (qJ), pitch angle (8) and yaw angle (l/J). These can be represented in the vector form n T = (qJ, 8, l/J) . Vehicle position is defined by the vector r T = (x, y, z). The body fixed frame is transformed to the inertial frame by the rotational matrix R where COSqJ is denoted by c<p and sinqJ is denoted by s<p: (C1/JceC1/Jses<p - s1/Jc<pc1/Jsec<p + S1/JS<P) R = \ufffd\ufffd\ufffd\ufffd\ufffd+\ufffd\ufffd\ufffd\ufffd\ufffd-\ufffd\ufffd -seces<pcec<p (1) The position of the quadrotor is obtained by Newton's second law of motion as: m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure9-1.png",
        "caption": "Fig. 9: Saturation map and current density map from the nonlinear MS simulation of the IM LR working point at full voltage and frequency.",
        "texts": [
            " Restrictions apply. 0093-9994 (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. current by this harmonic is sinusoidally distributed, along the rotor angular coordinate, for any level of saturation and rotor frequency. Thus, the hypothesis made in [14], about the spatial behavior of the rotor reaction is validated. The non-linear TH simulation in Fig. 9a and Fig. 4b represent the benchmark for the procedure described in the following, that have the scope to include the skin effect in the MS analyses of the IM. In the TH analyses, the rotor bars are modeled as plain conductors and no circuit property are imposed. A more complete TH simulation is reported for the final comparison, in which the whole machine is modeled and a cage circuit is defined to guarantee that the sum of the bars currents is zero. A further step for the MS analyses is including the effect of the actual rotor frequency to the rotor current: it can cause a non-uniform distribution",
            " As a manner of facts, considering the actual bar current distribution, the procedure in Fig. 3 would be supplied by a better estimation of the rotor parameters (resistance and leakage inductance) that can be derived from the field solution represented in Fig. 1a [14]. The consequence is also a better estimation of the rotor induced current for a certain slip and stator voltage supply. Carrying out the passages in Fig. 3, considering an uniform rotor current distribution, the field solution that satisfy the voltage condition is represented in Fig. 9. The procedure can be improved making a TH linear simulation to get the actual rotor bars current distribution. Both the stator and the rotor are supplied with equal and opposite currents in order to reproduce what happens along the q-axis. From the TH simulation field solution in Fig. 10a, the real part of the solution represents the time instant of interest to store the current distribution in each slots sectors. From the field solution in Fig. 10a the current in integrated in each slot regions and the ratio with the total bar current gives the part of the bar current that flows in a certain slot sector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002513_s11431-016-0561-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002513_s11431-016-0561-5-Figure1-1.png",
        "caption": "Figure 1 Schematic of the inclined-ellipse dimpled gas face seal and thermal boundary conditions.",
        "texts": [
            " In the present paper, the influence of the thermal effect on an inclined-ellipse dimpled gas face seal is analysed to investigate the temperature distribution and sealing performance considering thermal distortions at high rotational speeds of up to 35000 r min\u20131. The influence of the geometrical parameters of inclined dimple patterns on the temperature and pressure distributions is analysed. Then, a parametric study is conducted to quantify the changes in the sealing performance. An inclined-ellipse dimpled gas face seal consisting of a smooth rotor with rotational speed \u03c9 and a micro-dimpled stator with dimple depth hd is shown in Figure 1. As shown in Figure 1(a), the micro-dimpled region adjacent to the external radius ro with high pressure stretches from the dimpled radius rg to the external radius ro, and the ambient pressure is set at the internal radius ri. The balance radius rb is fabricated on the back side of the rotor, and gas flows along the direction of pressure decrease between the clearance h of two rings. Seal face distortions are taken into consideration while referring to the influence of the thermoelastohydrodynamic behaviour, and the film thickness of the lubrication region h is described as      d Deform, , , ,    h r c h r h r (1) where c is the clearance; r and  are the radial and circumferential coordinates; hd is the dimple depth; and hDeform is the clearance variation owing to face deformations. Inclined dimples are distributed uniformly around the centre axis of the rings. To define the percentage of the face area with dimples from rg to ro, a parameter called the area density of Sp is presented as follows: \u03b8 r p 2 2 o g ,  n n ab S r r (2) where n\u03b8 and nr are the dimple numbers in the circumference and radial directions, respectively. Inclined dimples are processed on the surfaces of the stator ring. An inclination angle \u03b2 is applied to describe the dimple orientation, as shown in Figure 1(b), which is regulated as the measuring angle from the long radius to the major axis along the gas flow direction. The long and short radii of the inclined dimples are denoted as a and b, respectively. The slender ratio \u03bb is used to describe the influence of the inclined-ellipse dimpled orientation, and it is expressed as .  a b (3) Therefore, the orientation of the inclined-ellipse dimple can be uniquely determined by the inclination angle \u03b2 and slender ratio \u03bb. Table 1 shows the dimensions of gas face seals with inclined-ellipse dimples",
            " It is summarized as follows. Table 1 Dimensions of the inclined-ellipse dimpled gas face seal Item Symbol Dimensions and data Outside radius ro 40 mm Inside radius ri 34 mm Dimpled radius rg 36 mm Balance radius rb 36.5 mm Area density Sp 0.25 Slender ratio \u03bb 3 Dimple numbers in circumference n\u03b8 150 Dimple numbers in radial direction nr 5 Ring thickness h1,h2 10 mm Considering the periodicity condition, a radial dimpled column can be used as the research object to treat the gas face seal, as shown in Figure 1(b). The Reynolds equation for ideal gas lubrication is  3 3 6 ,                            hh p h r p r r r r r (4) where  is the gas viscosity of N2 (nitrogen); in the present study, it varies with the compressibility, and it can be calculated based on [17].  is the gas density; and  is the rotational speed. Moreover, the gas pressure p is assumed to depend on the gas density and energy of gas molecules as p d m ,p c i E (5) where id is the degree of molecular freedom (5 for N2); Em is the energy of gas molecular per freedom",
            " The heat conduction equation for a stator is 2 2 s s s 2 2 2 0.              T T T r r r r r z (10) The heat conduction equation for a rotor with velocity along the \u03b8 direction is 2 2 c2 s s s s 2 2 2 s2 s2 ,                     k T T T T r c r r r r z (11) where s2 and cs2 represent the density and specific heat capacity of the rotor, respectively; Ts is the temperature of seal rings; and z is the axial coordinate. Heat boundary conditions including convective, adiabatic, and imposed ones are shown in Figure 1(a). The heat exchange condition from rings to ambient gas is  s c1 gs1 s1 0 s ,      T k k T T n (12a)  s c2 gs2 s2 0 s ,      T k k T T n (12b) where n is the direction perpendicular to the solid boundary; kc1 and kc2 represent the thermal conductivity of the stator and rotor, respectively; kgs1 and kgs2 represent the convection heat transfer coefficient at ambient boundaries of the stator and rotor, respectively; T0 is the ambient temperature. The heat transfer condition between the gas film and seal rings in the lubricating region satisfies the following expressions:  s c1 s1 s1 s ,      T k k T T n (13a)  s c2 s2 s2 s .      T k k T T n (13b) The convection heat transfer coefficients in eq. (13) can be obtained using 1 3 r c-gas s1 s2 3 ,  Q k k k c (14) where Qr is the Prandtl number, and it is given as p r c-gas ,   C Q k (15) where kc_gas is the gas thermal conductivity (0.024 W m\u20131 K\u20131 for N2), and Cp = 29.1 J mol\u20131 K\u20131 is the gas isobaric specific heat capacity. The pressure distributions and density of the gas film can be obtained using eq. (4) with a pressure-periodic boundary condition, as shown in Figure 1(b), as o o( , ) ( , ) , ( , / ) ( , / ).              , i ip r r p p r r p p r N p r N (16) where pi is the ambient pressure; po is the seal pressure; N is equal to the dimple number in the circumference direction. The temperature fields of the gas film can be obtained by eq. (7) with a dynamic temperature boundary o inlet r o( , ) , if ( , ) 0,    T r r T q r r (17a) i i r i( , ) , if ( , ) 0,    T r r T q r r (17b) where Ti is the temperature of the sealed gas, and Tinlet is the gas temperature at the inlet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002827_j.optlaseng.2020.106065-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002827_j.optlaseng.2020.106065-Figure14-1.png",
        "caption": "Fig. 14. The state of the particle and the surrounding fluids at 18 \ud835\udf07s: (a) where the powder particle is located; (b) speed distribution of the particle and the surrounding fluids; (c) the speed distribution of the pool around the particle.",
        "texts": [
            "03 m s \u2212 1 , hich is ~81.2% of its initial speed (combing with Fig. 13 (b), which hows the speed distribution of the particle and the surrounding fluids). ig. 13 (c) is the speed distribution of the melt pool around the partile, which demonstrates that the peak speed of the melt pool decreases o ~2.03 m s \u2212 1 , and the fluctuation area at the gas-liquid interface of he melt pool around the particle increases to more than twice the cross ection of the particle. When the powder particle moves to the position shown in Fig. 14 at moving time of 18 \ud835\udf07s and has an immersion depth about one-half of ts own diameter in the melt pool, the powder particle speed decreases o ~1.86 m s \u2212 1 , which is ~74.4% of its initial speed. As can be seen rom Fig. 14 (c), the fluctuation zone of the melt pool increases in depth urther, and the fluctuation area at the gas-liquid interface increases ignificantly compared to that at 12 \ud835\udf07s. The peak speed of the melt ool appears around the particle, which is ~1.86 m s \u2212 1 , and the speed radient was reduced compared to those at 6 \ud835\udf07s and 12 \ud835\udf07s. Fig. 15 shows the state of the powder particle and the melt pool at 9 \ud835\udf07s as the particle is immersed in the melt pool to a depth of twohirds of its diameter. It can be seen that the liquid level of the melt ool around the particle is lower than that of the mean level of the gasiquid interface ( Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003834_978-0-262-33936-0-ch053-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003834_978-0-262-33936-0-ch053-Figure1-1.png",
        "caption": "Figure 1: ARGoS simulation screenshot of the base and a deposit Dm away from the base edge. The base consists of a resting bay, an observation bay and an unloading bay. A light source is placed above the centre of the base to guide robot navigation. Pellets collected by the robots temporarily accumulate in the unloading bay, causing congestion.",
        "texts": [
            " If unloading is taking too long, a forager will tremble dance around the nest, inhibiting other bees from recruiting and thus reducing the number of foragers. Our social RTM uses a similar principle. All our experiments are performed in the ARGoS simulation environment which implements realistic 3D physics and robot models (Pinciroli et al., 2012). The simulation has continuous space and updates 10 times per simulated second. A circular base with a diameter of 3 metres is situated in the centre of the experimental arena. The base is divided into three sections (Figure 1): an interior circular resting bay with an annular observation bay around it and an annular unloading bay around that. A light source, placed above the centre of the base, is used by robots as a reference for navigation towards and away from the base centre (as in, e.g., Krieger and Billeter, 2000; Pini et al., 2013). Cylindrical resource deposits with radius rD are placed outside of the base, each containing an unlimited volume of resource. In order to enable robots close to a deposit to move towards it, a colour gradient with radius rC is present on the floor around each deposit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001391_icems.2013.6713350-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001391_icems.2013.6713350-Figure2-1.png",
        "caption": "Fig. 2. (a) Semi-closed slots. (b) Open slots with semi-magnetic wedges",
        "texts": [
            " This means that large 62nd harmonic could lead to large 4th force harmonic and subsequently high vibration level. In addition to open slots, the machine under investigation is also modeled with semi-closed slots and also with semimagnetic wedges. In the case of semi-closed slots, the slot opening is reduced by 50% compared to open slots. Two cases are modeled with semi-magnetic wedges, having \u00b5r=5 and \u00b5r=10 as relative permeability. The semi-magnetic wedges are inserted in the opening of the open slots. Semi-closed slots and open slots with semi-magnetic wedges are shown in Fig. 2. The main (58th) and 62nd spatial harmonics of radial component of flux density in the airgap as well as the main harmonic of tangential component (58th) are compared in Table I for four different cases. According to the table, open slots and semi-closed cases have the lowest and highest values of main harmonic in radial flux density. Using semi-magnetic wedges leads to higher value of radial flux density in the airgap compared to open slots. Results show that in the case of semi-magnetic wedge 2 (\u00b5r=10), the radial component of flux density is slightly larger compared to the case of semimagnetic wedge 1 (\u00b5r=5)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003889_gt2014-25223-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003889_gt2014-25223-Figure8-1.png",
        "caption": "FIGURE 8: Example of a modal shape measurement given by the SLDV",
        "texts": [
            " This modal analysis is conducted by a standard curve-fitting of frequency response functions (FRF) measured separately on the structures. Sample frequency ranges containing the initial family of modes are given in Figures 4 and 5. For confidentiality reasons, frequencies are normalized with respect to the mean frequency of the first family of modes of the two structures at rest. A modal analysis has been performed for the first four mode families on the impeller (Fig. 6) and for the two mode families on the casing (Fig. 7). Figure 8 illustrates 3 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use one modal shape of the casing measured along the axial direction using SLDV for one mode of the first family. The natural frequencies, damping factor and mode shapes, along with their nodal diameter numbers, have been extracted from measurement data. The modal analysis was completed under various thermal conditions and have revealed a decreasing frequency rate vs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003373_10667857.2020.1810926-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003373_10667857.2020.1810926-Figure10-1.png",
        "caption": "Figure 10. (a) Specimen\u2019s orientation in MagicsRP and (b) specimens fabricated (0\u00b0 orientation) on EOSINT P395 m/c.",
        "texts": [
            " Offsetting and thickening of Palmar side surface part. Step 3: additive manufacturing of splint Fabrication of specimens for strength testing. Prior to fabrication of wrist splint, mechanical properties of polyamide-PA2200 specimens are measured using tensile, flexural and compressive strength. Test specimens for tensile, flexural and compressive strength are prepared according to ISO 527\u20134, ISO 178:2010 and ISO 604:2002 standard respectively. Specimens are kept in xy-plane at three different orientations, i.e. 0\u00b0, 45\u00b0 and 90\u00b0 as shown in Figure 10(a). Three specimens per orientation for each test are fabricated. Specimens are fabricated on EOS P395 selective laser sintering machine (EOS GmbH, Germany). Other process parameters are chosen as follows: CO2 Laser power: 50 W, Laser beam diameter: 0.6 mm, Layer thickness: 0.12 mm, Build temperature: 175\u00b0C. Fabricated specimens are shown in Figure 10(b). Mean values of various strengths with error bars of one standard deviation are depicted in Figure 11. It is found that specimens fabricated along x-axis in xy-plane possess maximum tensile, flexural and compressive strength. So, it is decided that the wrist splint be fabricated along x-axis in xy-plane to obtain the best results. Fabrication of wrist splint. To fabricate the wrist splint, STL file of CAD model is loaded in MagicsRP for checking of errors viz. gaps, inverted normals, bad edges etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003271_s42243-020-00456-3-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003271_s42243-020-00456-3-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of sample position (a) and thermal deformation process (b)",
        "texts": [
            " TC18 plate with geometrical sizes of 400\u00a0mm \u00d7 40\u00a0mm \u00d7 80\u00a0mm was fabricated by LMD in a chamber with argon atmosphere (oxygen content less than 50 \u00d7 10\u22126). The scanning parameters used in the present study were laser beam diameter of 5\u20137\u00a0mm, laser power of 6\u20138\u00a0kW, scan speed of 800\u20131200\u00a0mm/min, and powder feeding rate of 800\u20131000\u00a0kg/h. After deposition, samples with dimension of \u03d510\u00a0mm \u00d7 15\u00a0mm were cut from LAMed (laser additive manufactured) components, and the sampling direction was as shown in Fig.\u00a01a. Then, the samples were heated to temperatures of 750, 800, 850 and 900\u00a0\u00b0C at a heating rate of 10\u00a0\u00b0C/s on a Gleeble 3800 thermal simulator, and uniaxial isothermal compression experiments were performed after 5\u00a0min of incubation. T\u03b2 transus temperature of TC18 alloy is around 890\u00a0\u00b0C [18]. During test, a tantalum alloy piece was introduced between samples and loading sector to reduce the friction and heat transfer as well as avoid bulging when the sample was deformed. The thermal compression process was carried out in a system using a set 60% deformation and strain rate of 1\u00a0s\u22121 at different temperatures. After deformation, samples were immediately water quenched (WQ). The detail of thermal compression process was sketched in Fig.\u00a01b. During thermal deformation, the real temperature, deformation time, force value and displacement of the samples were monitored and recorded by a digital system (Table\u00a01). Finally, all deformed samples were heat-treated according to the traditional TC18 titanium alloy double-annealing schedule. As shown in Fig.\u00a02, the process can be divided into three steps: First, sample was incubated at 830\u00a0\u00b0C for 2\u00a0h and furnace cooled (FC) to 750\u00a0\u00b0C holding for 2\u00a0h, and then air-cooled (AC) to room temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002476_tac.2016.2631302-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002476_tac.2016.2631302-Figure2-1.png",
        "caption": "Fig. 2: (a) Gershgorin discs with heterogeneous positive gains, (b) multiple negative gains, (c) two negative gains (0 < \u03b3 < 1).",
        "texts": [
            " If 0 < \u03c1 < 1 then all the eigenvalues, except the one at (1 + j0), are inside the unit circle, shown in Fig. 1b. This guarantees positional consensus of the system. This result is identical to the one obtained in [2] which uses Gershgorin\u2019s Theorem to establish stability for 0 < \u03c1 < 1. Here, the eigenvalues are explicitly given. The heterogeneous case is now considered. Lemma 1. Consider the system described by (1)-(2). If 0 < \u03c1i < 1, \u2200i, then the system achieves positional consensus. Proof. As the union of the Gershgorin discs lies inside the open unit circle in Fig. 2a (dotted circle), except at the point (1, 0) where the Gershgorin discs are tangential to the unit circle, it suffices to ensure that there are no repeated roots at 1. Observe that the characteristic equation for A, given by i=n\u220f i=1 (s\u2212 1 + \u03c1i)\u2212 i=n\u220f i=1 \u03c1i = 0 (3) does not have repeated roots at 1. This follows from the fact that the derivative of \u220fi=n i=1 (z+ \u03c1i)\u2212 \u220fi=n i=1 \u03c1i with respect to z at z = 0 is positive and is not zero if 0 < \u03c1i < 1, \u2200i. To obtain the point of convergence, the following summation (modulo n) is considered, i=n\u2211 i=1 zi(t+ 1)\u2212 zi(t) \u03c1i = i=n\u2211 i=1 (zi+1(t)\u2212 zi(t)) = 0 \u21d2 i=n\u2211 i=1 zi(t+ 1) \u03c1i = i=n\u2211 i=1 zi(t) \u03c1i , \u2200t",
            "3 of [7] the stability of the system in (1)- (2) can be examined, by setting the deviations of all the agents to zero. However, [7] only considers positive heterogeneous gains for the agents and thus the framework in [7] cannot be used to investigate the consensus of cyclic pursuit with one or more negative gains. To investigate whether the system given by (1)-(2) can tolerate negative gains and yet achieve consensus, consider (2). Suppose at least two gains \u03c1p and \u03c1q are negative. Then, two Gershgorin discs lie outside the unit circle (see Fig. 2b). Consider a parameter \u03b3 \u2208 [0, 1) and the matrix A\u0304 obtained by replacing the off-diagonal \u03c1p and \u03c1q in A (in (2)) with \u03b3\u03c1p and \u03b3\u03c1q , respectively. The Gershgorin circles are shown in Fig. 2c. From Gershgorin\u2019s theorem, it follows that A\u0304 has at least two disjoint discs, outside the unit circle, for \u03b3 < 1 (Fig. 2c) and therefore, at least two of its roots must lie outside the unit circle [11]. Since the eigenvalues of A\u0304 are continuous functions of \u03b3, as \u03b3 approaches unity, the two roots may migrate towards the unit circle, but they must belong to the union of the discs. So they can only enter the unit circle through the point (1, 0). However, for \u03b3 < 1, there are two roots outside the unit circle. Hence, due to continuity of root locus, at \u03b3 = 1, the roots may at most migrate to the point (1, 0) resulting in multiple roots at 1, thereby ruling out consensus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003708_5.0015728-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003708_5.0015728-Figure16-1.png",
        "caption": "FIGURE 16. Cycle Fatigue Simulation Results",
        "texts": [],
        "surrounding_texts": [
            "040011-7 Based on figure 19 is an S-N diagram graph that occurs in AISI 4315 material. The maximum cycle value at the 1.5 mm crack distance is 87.04 with a stress of 1498.6 MPa, compared to the distance between the 2.5 mm crack which has a maximum stress value of 1408 MPa and a cycle of 100.89. Whereas the connecting rod without crack has the lowest value, which is 886.9 MPa with a cycle of 325.91. Based on the results that have been obtained, it appears that the stress will be even greater in the specimen that is cracked, especially in the crack area close to the big end distance of 1.5 mm. The farther the crack from the big end, the less alternating stress value is obtained [169]. Fatigue limit obtained in the simulation results is 1E + 06, according to the statement that the fatigue limit value that occurs in steel is 1E + 06 [17,18], so it can be seen that the cause of failure on the component is not due to lack of material strength, but rather as a result of cracks due to repetitive cycle loads, as well as more loads that trigger cracks near the big end. 040011-8"
        ]
    },
    {
        "image_filename": "designv11_30_0002533_iros.2016.7759198-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002533_iros.2016.7759198-Figure3-1.png",
        "caption": "Fig. 3. Dynamic model of tracked vehicle colliding with a wall",
        "texts": [
            " Then the tracked vehicle cannot leave the wall and moves in the opposite direction. This phenomenon is observed when the robot moves through narrow passages during a plant inspection. In these cases, the tracked vehicle cannot move according to the control commands and cannot continue the inspection. The above problems can be solved by using a motion control method based on the contact or collision force. We derive control rules that allow the tracked vehicle to move according to control commands. A dynamic model of a tracked vehicle is shown in Fig.3. Table I shows parameters of the dynamic model. This model considers the collision force. A coordinate frame is attached to the center of the vehicle body, whose x-axis faces front and y-axis faces left. Below, the model is analyzed on the frame. Although this model considers collisions in the rear right corner, analysis is also available for all other corners by inverting the l1 or l2 parameters. In each case, the dynamics of the model are the same because the model is just inverted geometrically",
            "(1)-(3), TABLE I PARAMETERS IN DYNAMIC MODEL F1 driving force of right track F2 driving force of left track F3 slide fliction force of right track F4 slide fliction force of left track F5 nomal force from a wall F6 friction foce from a wall 2d tred \u03b8 angle between wall and robot l1 distance between center of gravity and collision point in Y-axes l2 distance between center of gravity and collision point in X-axes (Lx, Ly) location of rotation center on robot coordinate system \u00b5g maximum coefficient of static friction between track and ground vr peripheral velocity of right track vl peripheral velocity of left track t time for constant acceleration respectively: Mx\u0308 = F1 + F2 + F5 sin \u03b8 + F6 cos \u03b8 (1) My\u0308 = \u2212F3 \u2212 F4 + F5 cos \u03b8 \u2212 F6 sin \u03b8 (2) J\u03b8\u0308 = F1(Ly + d) + F2(Ly \u2212 d)\u2212 (F3 + F4)Lx +F6 \u221a (Ly + l1)2 + (Lx + l2)2 (3) The straight line passing through rotation center \u201dO\u201d to collision point \u201dA\u201d is vertical to a wall. This is because point \u201dA\u201d has a geometrical constraint in which it is bound to move on the wall surface. Based on this constraint, the relation between Lx and Ly is described in Eq.(4): l2 + Lx l1 + Ly = tan \u03b8 (4) Using the dynamic model shown in Fig.3, the motion of the tracked vehicle can be analyzed when its side surface or tracked surface collides with a wall. In this chapter, an analysis of a side-surface collision is conducted. An analysis of a tracked-surface collision is conducted in section B in chapter five. The model shown in Fig.3 describes a simple case of collides at just one point. However, it can alseo be used for collision between the side surface of the robot and a wall. In this case, the angle between robot and the wall (\u03b8) is zero. For a linear motion, the passive wheels absorve the friction with the wall and a collision force is not generated. As the robot departs from a wall, a collison forcece is generated at just one point due to the circular motion inherent in leaving the wall. Analysis is conducted by substituting some conditions for the equations of motion",
            " \u03c9ref =  \u03c9 (if no collison) \u03c9 (else if |\u03c9| < \u03c9max) sig(\u03c9) \u00b7 \u03c9max (else) (17) vref =  v (if no collision) v +Kp \u00b7 (v \u2212 vc) +Ki \u00b7 \u222b (v \u2212 vc)dt (else) (18) Here, \u03c9max is a limitation of angular velocity, sig() is a function which returns sign (plus or minus) of an input value, and vc is current linear velocity.The collision is detected by using differencd of the angular velocity(\u03c9ref \u2212 \u03c9c). When a tracked vehicle move in narrow passages, its front side sometimes collides with walls and is lifted up. On dynamic model in Fig.3, F6 cannot be ignored at collision. In that case, if vehicle moves in backward, F6 generates moment around rotation center. Then, moreover, the vehicle rotates and climbs the wall. To prevent the above undesirable motion, driving force F1 and F2 should be controlled by filling a condition that \u03b8\u0308 < 0 in Fig.3. When F6 is predominant in the rotational motion, F3 and F4 do not affect the rotational motion. Therefore, F3 and F4 can be considered as 0. Eq.(19) is derived by substituting those conditions into Eq.(3). s 0 < F1(Ly + d) + F2(Ly \u2212 d) +F6 \u221a (Ly + l1)2(1 + tan \u03b8) (19) In Eq.(19), rotation radius Ly changes depends on F6. Here Eq.(20) is derived from Eq.(19) with a condition,Ly = \u2212l1, assuming that rotation center \u201dO\u201d and collision point \u201dA\u201d is equivalent in Fig.3 when the front part of vehicle is lifted up onto a wall. 0 > F1(l1 \u2212 d) + F2(l1 + d) (20) The problem that a vehicle lifting up a wall can be solved by deciding F1 and F2 using Eq.(20) when its tracked surface collides with a wall. In experiments bellow, collision on tracked surface is avoided by rotating counter clock wise when right track collides, or rotating opposite side in case of left track. In both cases driving force should be F2 = \u2212F1, which meets Eq.(20) in any angular velocity. For the experiments bellow, angular velocity is commanded in 45deg/sec when collision occured on tracked surface of sub-trackes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001504_amm.459.449-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001504_amm.459.449-Figure2-1.png",
        "caption": "Fig. 2 FE Meshed Model",
        "texts": [
            " Nonlinear material constants of SILASTIC P1, Silicone RTV based on Ogden model, presented in [6] were used for the current finger model. Table 1 shows the material properties and geometrical properties used in the analysis. The general form of Ogden model is, W \u03bb , \u03bb , \u03bb = \u2211 \u03bb + \u03bb + \u03bb \u2212 3 . (1) where, , , is strain energy function, and are material constants, , and are stretch ratios. Under the assumption of incompressibility, the Eq.1 can be rewritten as, W \u03bb , \u03bb = \u2211 \u03bb + \u03bb + \u03bb \u03bb \u2212 3 . (2) Meshing was done with 3D tetrahedron Solid 187 elements with \u2018patch conforming\u2019 option. Fig.2 shows the meshed FE model of the finger. Modes are inherent properties of a structure, and are determined by the material properties and boundary conditions of the structure. Each mode is defined by a natural frequency, modal damping and a mode shape which are so called modal parameters. In the current analysis, one end of the finger was constrained in all three directions as a boundary condition. An n-degree of freedom system may be described by the matrix equation of motion below; + + = . (3) Where M, D and K are the mass, damping and stiffness \u00d7 matrix, respectively; is the \u00d7 1vector of degrees of freedom and is the \u00d7 1 vector of forces applied to the structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002085_j.wear.2016.04.027-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002085_j.wear.2016.04.027-Figure5-1.png",
        "caption": "Fig. 5. Contact stresses and definition of coordinate system employed in the RCF analyses.",
        "texts": [
            " Under full slip conditions f equals the coefficient of friction \u03bc, and the interfacial shear stress is obtained as \u03c4xz \u00bc f Upz x; y\u00f0 \u00de \u00f08\u00de Here pz is the contact pressure, f \u00bc |Fx/Fz| the traction coefficient with F being resultant forces and x and z coordinates along and into the contacting bodies, respectively. Two cases of full slip (\u03bc\u00bc f\u00bc0.3 and \u03bc\u00bc f\u00bc0.35) are investigated. Under partial slip conditions the Carter solution [11] is employed strip-wise over the elliptical contact patch (cf [1]). Partial slip analyses are carried out for the case of a traction coefficient of f\u00bc0.3 and friction coefficients of \u03bc\u00bc0.4 and \u03bc\u00bc0.5. In Fig. 4 the interfacial stresses (at y\u00bc0 with coordinates following Fig. 5) under these conditions are visualized. A significant difference between peak interfacial shear stresses in the case of full and partial slip (for the same traction coefficients) may be noted. In shakedown based approaches to surface initiated RCF analyses this has been accounted for by employing the maximum value of the interfacial shear stress, \u03c4xz,max, see e.g. [12]. To investigate the validity of such an approach, stresses below the contact are evaluated using a procedure outlined in [13] that employs a numerical integration of contributions from a large number of point forces approximating the distributed surface stress",
            "3 548 1.40 292 1.09 276 1.11 \u03bc \u00bc 0.4, f \u00bc0.3 505 1.29 285 1.06 269 1.08 \u03bc\u00bc f \u00bc0.3 391 1 269 1 249 1 \u03bc\u00bc f \u00bc0.35 452 1.16 309 1.15 289 1.16 The Dang Van equivalent stress according to Eq. (4) is employed to characterize the severely multiaxial stress beneath the contact. In the current study, an optimisation procedure is employed to derive the mid value of the deviatoric stress tensor, see [8]. The equivalent stress \u03c3dv is evaluated at varying depths (zcoordinates) in the symmetry plane (y\u00bc0), see Fig. 5. Evaluated \u03c3dv-magnitudes for the studied cases are compiled in Fig. 6. It could be noted that at a depth of some 3 mm the influence of shear stresses applied at the surface is very low. To investigate gradient effects, the attention is focused to magnitudes of \u03c3dv at the uppermost 0.1 mm below the surface \u2013 Please cite this article as: A. Ekberg, et al., Stress gradient effects in s (2016), http://dx.doi.org/10.1016/j.wear.2016.04.027i i.e. in the region where cracks are presumed to initiate. Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000882_0954405414564405-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000882_0954405414564405-Figure11-1.png",
        "caption": "Figure 11. Diagonal hobbing method.",
        "texts": [],
        "surrounding_texts": [
            "Pressure angle a ( ) 20 Pressure angle a ( ) 20 Number of teeth ZC 15 Number of hob threads ZB 1 Normal module mn (mm) 6 Normal module mn (mm) 6 Total tooth depth h (mm) 6 Hob spiral angle n ( ) 10 Tooth width B (mm) 30 Spiral angle direction Right hand Outside diameter of gear blank DC (mm) 120 Outside diameter of hob DB (mm) 20 Gear spiral angle b ( ) 30 Spiral angle direction Left hand Installation height H (mm) 20 at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from workpiece. Then, the EGB was opened by the G81 command, and the hob was gradually fed to the full tooth depth along the X-axis. After a short pause, the hob was fed along the Z-axis over the entire tooth width, as the hob cut along the X-axis and Z-axis. Finally, the EGB was closed by the G80 command. During the cutting process, the single-axis tracking error of the X-axis and Z-axis was both less than 0.4mm, and the tracking error of the C-axis was less than 0.0006 rad. The pitch error of the workpiece reflects the accuracy of EGB and is illustrated in Figure 15, with a maximum absolute error less than 15mm. Diagonal hobbing method. In this section, the diagonal hobbing method is evaluated. In this case, the spindle speed was 900 r/min. The reference and actual trajectories are illustrated in Figure 16. The C-axis followed the B-axis, Y-axis and Z-axis throughout the machining process, and its reference position and actual position are illustrated in Figure 17. There is also an instantaneous position change at the moment when the EGB is opened, as the C-axis attempts to keep up with the spindle. The tracking error during the diagonal hobbing motion process is shown in Figure 18. It can be seen that the single-axis tracking errors of the X-axis, Y-axis and Z-axis are all less than 0.4mm during the hobbing process, and the tracking error of the C-axis is less than 0.0006 rad. The estimate of the gear pitch error in the at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from gear diagonal hobbing process is illustrated in Figure 19, and the maximum absolute error is less than 18mm. From Figures 12\u201319, we can conclude that the hobbing CNC and the EGB can work correctly for the examples tested. The C-axis follows the rotation speed of the spindle and the feed rate of the other servo shaft. From Figures 14 and 18, we can see that the tracking error of each axis is very small. Here, we use the gear pitch error as the performance metric of the EGB. As shown in Figures 15 and 19, under the control of the EGB, the maximum gear pitch error is 0.0147mm in the axial hobbing movement process, and the maximum at Purdue University Libraries on July 12, 2015pib.sagepub.comDownloaded from gear pitch error is 0.0176mm in the diagonal hobbing movement process. The results show that the proposed EGB is effective."
        ]
    },
    {
        "image_filename": "designv11_30_0001107_1.4023226-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001107_1.4023226-Figure1-1.png",
        "caption": "Fig. 1 Tilt-back ball screw machine and its control and cooling systems",
        "texts": [
            " The skewness parameter is applied in this study to deal with the number distribution of the peaks with various G and Ds values. The skewness value can provide the evolutional trend information of the vibration quantity and magnitude at different testing times. In the present study, a commercial tilt-back ball skew machine was provided as the test subject for the short- or long- range tribological and vibrational measurements. It can be set at 3000 rpm as the maximum rotational speed and with an acceleration of 0.6 g. This machine, as shown in Fig. 1, is comprised of a control box to monitor the rotational speed and acceleration of the motor, an oil cooling system which includes an oil pump and the central hollow oil channel in the screw, a water cooling system for motor cooling, a servomotor, a coupling for torque transmission, two ball bearing housings at the two ends of the ball screw, and a couple of linear guiders acting as the sliding supporter of the carriage. The specifications of the ball screw are shown in Table 1. The end cap return ball nut is designed as the ball recirculating mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002390_gt2016-57410-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002390_gt2016-57410-Figure14-1.png",
        "caption": "Figure 14. Tilting pad journal bearing with ISFD",
        "texts": [
            " The bearing manufacturer have references for the use of ISFD technology to solve field instability problems; the steam turbine manufacturer didn\u2019t have site experience with an ISFD bearing but had previously carried out an experimental activity with this kind of technology. In [17] the testing activity for the dynamic coefficients identification is presented, whilst in [18] a comparison of the results obtained during the test of a dummy rotor with and without ISFDs is reported, demonstrating the high damping capability of the device. SOLVING INSTABILITY PROBLEM An Integral squeeze film damper (ISFD\u00ae) is an SFD and spring that are integrated in one piece, so that it can be easily integrated with the bearing, as shown in Figure 14. The ISFD technology allows decoupling stiffness and damping coefficients of the bearing support. The segmented design of the damper prevents circumferential flows, often cause of cavitation, and allows higher damping over a wider range of operating conditions. Technical reviews on ISFDs and their principles are well described by Ertas et al [8]. An optimization of the stiffness and damping coefficients of the ISFD in the specific rotor-bearing system was conducted to find the optimum damping value at the full load conditions, including the effects of the seals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001724_iros.2015.7353938-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001724_iros.2015.7353938-Figure8-1.png",
        "caption": "Fig. 8. Snapshots of the robot stepping on a soft floor.",
        "texts": [
            " The computation time for the proposed control algorithm for this experiment is presented in Figure 7. It can be seen in Figure 7 that for the iCub robot with 38 DoF performing 13 motion and force tasks, and with a total task dimension of 68, the computation time for the control algorithm is within 10ms without any specific code optimization. Real time implementation of the proposed approach on a torque controlled humanoid robot can thus be envisioned. In this experiment, the robot keeps switching its stance foot on a soft floor (see Fig. 8). The soft floor is modeled as two separate movable planks, one under each foot. The controlled tasks include the CoM task, the moving foot task, the posture task, and the stance foot contact task. Each foot contact force is constrained to lie inside a friction cone to avoid foot slippage. As mentioned in section III-B, instead of manually switching the CoM target position between above the two feet, it is computed automatically based on the desired foot contact force by solving (7). The resulting foot contact forces and the profile of the CoM position is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000931_j.medengphy.2014.11.006-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000931_j.medengphy.2014.11.006-Figure10-1.png",
        "caption": "Fig. 10. Predicted deformations of soft tissue and skin in B\u2013C slice of shoulder-foam contact model: (a) before loading and (b) under 12 mm compression.",
        "texts": [
            "16 kPa in the bony prominence region. The peak von-Mises stress 3.16 kPa) on the surface is about 40% and 56% of the peak von-Mises tress at points B (8.01 kPa) and C (5.67 kPa), respectively. In Fig. 9(b), he high hydrostatic stress also occurs at the bony prominence reion with a peak value of 3.53 kPa (26.5 mmHg) at point D, where the kin first contacts the foam during compression. The peak hydrostatic tress is lower than the 32 mmHg threshold for PUs. .2. Deformation of the skin and soft tissue Fig. 10 shows the deformation of skin and soft tissues in the \u2013C cross-section before and after 12 mm compression. At point , the thickness of soft tissue in the Z direction is designated as hT B =13.1 mm). The thickness of skin hS B is 1.95 mm. At point C, the hickness of soft tissue hT C is 11.36 mm, while the thickness of skin S C is 1.80 mm. After compression, at point B, the thickness of hT B is educed to hT B\u2032 (=11.5 mm), about 12.2% strain. The thickness of the kin is reduced to 1.73 mm, about 11",
            " Personalized PU care can e achieved via better understanding of an individual\u2019s anatomical eometry, particularly the bone prominences with high stress conentration. Although, the sliced distance of the CT images (0.6 mm) nd pixel spacing (0.938 mm) in this study limited the accuracy of the tudy of bone prominence effect. The CT data with higher resolution ill be required in the future to study the radii of curvature of 3D one prominences and their effect on the stress concentration. For the deformation of skin and soft tissue in shoulder, three key bservations are obtained. As in Fig. 10, the deformation of the skin nd soft tissue is not uniform. Deformation of the skin is smaller han that in soft tissues because the skin has a higher Young\u2019s modlus. This finding confirmed the predictions conducted by Makhsous t al. [12,13], although the deformations of skin and soft tissue at ess and close-up views of the bone prominence at (b) point B and (c) at point C. s c ( ( v f m i t m a s a e p p p e a a o p t p T h s s a ( t 2 t c l a s s f s p s f w S t d a m a P P t p o a C F t 1 S i R [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ [ houlder region are smaller than that at buttock due to the anatomical onfiguration and loading conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002870_s12206-020-0231-3-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002870_s12206-020-0231-3-Figure1-1.png",
        "caption": "Fig. 1. Structural diagram of the 6-DOF parallel robot.",
        "texts": [
            " 2 introduces the 6-DOF high-speed parallel robot and builds the kinematic and dynamic model. Sec. 3 elaborates the establishment process of the comprehensive evaluation index. Sec. 4 discretizes the robot motion space and establishes the robot motion constraint conditions. Sec. 5 provides an example to illustrate the effectiveness of the proposed method. Sec. 6 describes the servo motor selection process. Sec. 7 summarizes the paper and discusses the prospect of future work. 2. Kinematic and rigid body dynamic analy- sis The 6-DOF parallel robot is shown in Fig. 1, and it consists of a static platform, a moving platform, and six identical independent kinematic chains [16]. Each kinematic chain uses the RSS (R for revolute joint, S for spherical joint) construction, which has a motor, a reducer, a master arm, a slave arm, and two spherical joints. The master arms produce rotation movement driven by the motor. The slave arm is driven by the spherical joint mounted at the end of the active arm. The other end of the slave arm is connected with the moving platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003488_icra40945.2020.9196838-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003488_icra40945.2020.9196838-Figure2-1.png",
        "caption": "Fig. 2: Inertial and body reference frames",
        "texts": [
            " As is common in UAVs, the North-East-Down (NED) convention is adopted such that the i\u03021, i\u03022, i\u03023 basis vectors of the inertial reference frame, F i, align with the North, East, and Down directions, respectively. Similarly, the body reference frame, Fb is affixed to aircraft centre of mass, and is selected such that the b\u03021 basis vector points out of the aircraft nose, the b\u03022 vector points out of the starboard wing, and the b\u03023 vector points out of the aircraft belly. These reference frames are shown in Fig. 2. The equations of motion of the UAV are p\u0307i = CTbivb (1) mv\u0307b = \u2212\u03c9\u00d7b mvb +mgCbik\u03023 + T k\u03021 + Faero + F\u03b4 (2) C\u0307bi = \u2212\u03c9\u00d7b Cbi (3) J\u03c9\u0307b = (J\u03c9b)\u00d7\u03c9b + M\u03b4 + Maero + Mgyro (4) where the UAV mass, m, and the second moment of mass, J, are considered constant since the body is considered rigid. The operator (\u00b7)\u00d7 is the skew symmetric operator as defined in [11]. The unit vectors k\u03021 = [1 0 0]T and k\u03023 = [0 0 1]T are defined for simplicity and will be used in the control derivation. The state space is composed by pi, the position of the centre of mass of the UAV in inertial frame coordinates, vb, the velocity of the centre of mass resolved in body frame coordinates, \u03c9b, the angular velocity of the body frame with respect to the inertial frame in body frame coordinates, and Cbi, the Direction Cosine Matrix (DCM) from inertial to body frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001849_iecon.2015.7392315-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001849_iecon.2015.7392315-Figure3-1.png",
        "caption": "Figure 3: Magnet and hall sensor installation.",
        "texts": [],
        "surrounding_texts": [
            "2\nto realize and validate with theoretical basis and technical support.\nII. AIRFRAME DESIGN Strategy of airframe design is to make two systems work without affecting each other in system stabilization. The minimal system for multi-rotor is the quad-rotor configuration. It consists of 4 propellers where 2 rotate counterclockwise and other rotate clockwise, to balance their rotation inertia. Each propeller has an adequate distance from its center of gravity and spread away from each other. The moments should be enough to form a stabilization in roll and pitch; while the yaw control is provided by differentiate rotational speed of counterclockwise and clockwise propellers. These are the principle of stabilizing the angular movement of quad-rotor system [6].\nThe design of current hybrid propulsion UAV is complicated. Because it requires such converter or generator to convert mechanical energy to electrical energy or vice versa [7]. The proposed QiQ design put the advantage of each kind of power plant right it their suitable position. Electrical motors provide stabilization because of advantage in response while the major lifting power is provided by gas engine. The disadvantage of two power systems are covered by each other\u2019s advantage which provide a perfect concept of propulsion system.\nFor simplification and cost effective in applications, symmetrical configuration of motor characteristics and position is used. This simplifies the stabilization controller as motorpropeller sets would ideally have the same characteristics and powered under the same control situation. Symmetrical configuration reduces complicate controller design as the motion of each axis are ideally decoupled. In the proposed QiQ system, the gas-electric dual power design is inevitable to use at least two kinds of propeller-motor sets. Cancelling torque generated by engine is preferred by the same kind of engine installation. Two from four engines should be counter rotating propeller in one system.\nThe proposed QiQ design considers more about the balancing power during flight. Instead of installing every energy source such as batteries or gasoline in the middle of the airframe like the small multi-rotor, this proposed QiQ design considers\nthe possibility to install energy source closer to the motor or engine. This may reduce longer wire losses in the electric power supplies. Also, longer fuel pipe may induce head loss from fuel tank to the engine too. But head loss in fuel flow is not as serious as the loss in electrical system. For small design this effect can be ignored, while in the proposed QiQ is, it becomes to larger size, these losses might become significant to pay attention to. Installing energy source away from center of the aircraft also increase space in the middle for other payloads. During flights, battery weight will maintain unchanged, but gasoline fuel changes in a great deal. The design consideration places the battery closer to its motor, but keeps the fuel tank the middle of the airframe.\nOne major drawback for separating energy sources is their consumption rate being not equal due to many factors such as payload installation, flight scenario, weight balancing or imperfection of the drive systems. The unbalance in battery voltages may affect the response and performance of motors as well as flight performance. Unbalancing energy consumption may accumulate over a long period of time. This will result in unstable flight performance or catastrophic situation if one battery run out of energy during flight. Four engine consideration instead of two will allow the controller to possibly compensate this unbalancing problem during flights. This capability has to be included in the development phase.\nIII. ENGINE CONTROLLER DESIGN Multi-rotor requires accurate RPM response from each actuator. Gas engine power output is usually controlled by the air inlet value, and its performance is complicatedly related to outside environment such as pressure and temperature. But engine performance can be optimized by adjusting fuel-air mixture. A common design of engine ignition has two needles where one is effective to lower RPM and the other is effective for higher RPM. Adjusting these two needles may change engine characteristics and performance accordingly. Engine temperature also plays an important role to engine performance. These factors are major reason that would not allow open loop control for RPM controller as the environment different might induce characteristic variation and affect the flight performance dramatically.\nEngine carburetor design is an important factor to RPM controller. In the proposed QiQ design, GT-33 from O.S. Engine is adopted [8]. Datasheet from manufacturer stated that 70-80% of maximum engine power develops with half throttle. From pilot experiment also shows that engine response to throttle input is nonlinear to the throttle angle. So that the linearization of throttle servo to RPM output need to be normalized first. Then the incremental PD controller can be implemented to control RPM of engine. These two combinations allow the system to response to instant movement of RPM command as well as accurately tracking the optimum servo position for RPM control even there are environment factors.\nFlight controller core is running on STM32F427 microcontroller from ST electronics [9]. The engine controller core is running on STM32F103. The flight controller handles stabilization of the multi-rotor as well as communication to the ground center. Attitude commands are sent from R/C transmitter to the R/C receiver and then to the flight controller. Flight\n001514",
            "3\ncontroller converts commands as the input of controllers. Then it sends actuator commands to brushless electronics speed controller and engine controller. There are two communication channels between flight controller and engine controller to handle discrete command and continuous command. On engine ignition and shutdown sequence, engine controller will command engine upon CAN Bus message. Once everything is stable, engine controller will listen to the streaming Pulse Width Modulation (PWM) instead for continuous update. The control system schematic is as shown below.\nEngine controller senses rotation speed from hall sensor installed on the engine. The same hall sensor is also used for engine\u2019s Capacitor Discharge Ignition unit (CDI). Engine speed data is updated every engine turn by using the time difference from each received pulse.\nControl loop of the engine controller is running at 50Hz update rate which matches the servo motor input frequency. Incremental PID controller is used for this design. The desired engine rotational speed (in RPM) is called setpoint. The difference between measured RPM and the setpoint is the error (E). The incremental PID controller\u2019s control algorithm is as follows:\n(1)\nwhere k is the sample number (k=0, 1, 2 \u2026), U(k) is the output at k iteration, is RPM errors at iteration k, and are proportional gain, integral gain and derivative gain respectively. According to formula above, can be obtained from recursive method:\n(2)\nThen, we can calculate the incremental step from subtraction of (1) and (2) above:\n(3)\nwhere is the incremental output and it will be added to the last output which is . This is the incremental PID algorithm as used. is the key part of the engine controller as it allow the incremental controller to response to error accordingly. is required to response to immediate error and fast change of setpoint. is used for damping the oscillation and allows larger for faster response without overshooting.\nBecause of the fact that engine response to throttle input is not linear, the use of only a set of controller parameter is not suitable for whole system. From carburetor design, the engine response curve is steep at low throttle. This becomes less sensitive at higher throttle. Actuator linearization is used as a normalizing the response rate of the engine. This is done by capturing relation between RPM and throttle position along full range of engine at operation condition. Data is then cut in to two sessions to form two equations for each section of throttle.\nIV. EXPERIMENT RESULTS AND ANALYSIS In the QiQ airframe design and fabrication, gasoline engine model GT-33 from O.S. Engine is adopted into experiments. It was equipped with 16 inch diameter and 10 inch pitch 3-blades propeller from Master Airscrew, oil-gas mixture at 3.5% by volume, under ambient temperature at 27 . An infrared noncontact thermometer MLX90614 from Melexis [10] is used to measure temperature of exhaust pipe of engine at the point where exhaust gas outlet from engine is connected. Although measuring temperature from exhaust nozzle might not represent the internal temperature but it can indicate the overall temperature of the engine during test. The sensors has accuracy of when the engine is at the highest exhaust temperature and when engine is at the lowest engine temperature.\nBefore implementing controller to engine, the controller linearization needs to be set. This is done by start and warm up engine until temperature is saturated and stable. Then increase throttle slowly while monitoring throttle and engine RPM. Since the stability of engine under 4000 RPM is poor and engine may subject to cool down. The controller is designed to work at the region higher that this point.\n001515",
            "4\nFrom Figure 4, it can be seen that the relation between engine rotating speed and throttle input is highly nonlinear. If the PID controller is implemented directly to the throttle, the tuned PID parameter may not be optimized for all operation range. The linearization is done by using 3-point spline where the middle point data is from the throttle at 6000RPM. Controller output is the scaled first which makes the system more linearized.\nLong term stability is important in the proposed QiQ multirotor flight controller system. Controller should have accurate and similar characteristic even there is environment disturbance. Temperature is one of major factors in engine characteristics. Controller should be able to compensate engine response while it has not reached the design temperature and restored its compensation fast enough. The engines were conducted experiments to observe engine characteristics and performance without controller. An engine is start from cool down state then was subjected to immediate throttle input. Then the throttle input was reduced to observe the cool down response. The response is as shown in Figure 5.\nAt ignition moment (t=0), engine RPM ramped up sharply, as shown in Figure 5(a). The RPM, however, is hesitate to increase to stable region for few seconds. This is because of the temperature of engine is not high enough to operate at condition. Figure 5(b) shows the exhaust pipe temperature rose sharply at the beginning. Although exhaust pipe temperature only show trends of overall engine temperature but it still indicates the engine condition at that moment. Engine RPM is stable after 15 seconds which is unacceptable to be implemented directly into the proposed QiQ multi-rotor system. Large asymmetrical thrust might greatly affect flight stability and performance.\nAt t=90s, engine throttle was reduced from 35% to 17%. RPM drops accordingly. After a period of time while engine was cooling down, engine RPM continues to drop from 4100 to 3720 in a minute. This shows the long term instability of open loop system which requires compensation by controller to enable stability for the whole flight.\nEngine RPM control experiment using incremental PID has been implemented. The engine was started from cool down state. The controller was commanded to increase the set point from 4750RPM to 6700RPM by pilot\u2019s controller knob. After 75 seconds, setpoint was changed to 4960RPM and 4350RPM after 60 seconds. Controller setpoint was once again raised to 6700RPM for 60seconds and then returned to 4350RPM.\nFrom Figure 6(a), the engine RPM controller managed to track the commanded setpoint accurately even at the sharp edge and instantaneous change. Engine RPM output is controlled and has similar characteristics even the engine temperature during overall test period was changed by accumulated heat. The effect of controlled is clearly seen from Figure 6(b) where controller to compensate throttle requires to sustain engine RPM at that moment when the engine temperature is not high enough. At t=60s, throttle input to engine is 69% which is later reduced to 46% after 26seconds. This throttle compensation can be seen again at t=260s. It is obvious that the controller compensation is required to allow engine RPM stay stable at the desired point during transient change of setpoint. Controller performance at steady state can be seen from t=290s where the engine RPM has no obvious change but throttle output has already responded to the small fluctuated error. At lower RPM from t=95 to t=260 and from t=318 to t=400, the engine temperature drops which make the engine RPM output lower at the same throttle input. But engine RPM controller slightly compensates it by increasing throttle slowly according to temperature drop. This proves the effectiveness of controller in both high and low temperature.\nAnother experiment is also carried before flight test. This experiment is to test controllers\u2019 characteristics before flight. It\u2019s important to assure the similarity of the thrust actuator for the proposed QiQ multi-rotor system. Unbalance thrust may generate excess moment which destabilizes the QiQ. This experiment completes igniting all engines and enables RPM controller. After that, step inputs are given to see their responses.\n001516"
        ]
    },
    {
        "image_filename": "designv11_30_0000918_icems.2013.6754383-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000918_icems.2013.6754383-Figure1-1.png",
        "caption": "Fig. 1. Proposed motor.",
        "texts": [
            " Torque, output, mechanical strength of the rotor, loss, efficiency, and demagnetization are discussed in this paper. The motor proposed for small EV is aimed to attain the maximum output of 15 kW and be able to drive at the maximum speed of 10,000 rpm. II. PROPOSED MOTOR Table I shows the specifications of the motor proposed. The permanent magnet, NMF-12E, is the ferrite magnet, and the stator has distributed windings. Size and rated values of the motor proposed are the same as a motor with bar rare-earth magnets for 15 kW. Fig. 1 shows the motor proposed. This motor produces higher reluctance torque by arranging magnets along the magnetic flux flow [4]. It consists of three magnet layers [5], and there are ribs in the third layer. The width of the rib is 0.5 mm. III. RESULTS OF ANALYSIS The motor proposed with ribs and the motor Type A with no ribs in the third layer are shown in Fig. 2(a) and (b), respectively. The characteristics of the motor proposed are compared with Type A to demonstrate the effect of the ribs. 978-1-4799-1447-0/13/$31",
            " posed is more than 90 % e maximum efficiency is % because the output of posed. E. Demagnetization Analysis As is well known, the ferrite magnet is low temperature. Fig. 13 shows the result of analysis at -40 \u00b0C. The analysis condition i current is 1.5 times of the rated current (the m and the phase angle \u03b2 is 90 degrees [7]. Th density of the irreversible demagnetization -40 \u00b0C. The irreversible demagnetization is the ribs in the third layer of magnets as shown is not observed in Type A as shown in Fig. 1 analysis condition as the motor proposed. According to results of analysis, the ma this motor decreases by 0.7 %, and is 55.7 irreversible demagnetization begins as shown IV. CONCLUSION This paper proposed the motor which can high output and the high speed drive. Simula that the motor proposed achieves 100.7 % o maximum output of 15 kW and 109.0 % f speed of 10,000 rpm. Loss, efficiency, and demagnetization discussed in this paper. The efficiency of the is more than 90 % in the wide speed range"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003484_0954406220963148-Figure19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003484_0954406220963148-Figure19-1.png",
        "caption": "Figure 19. Path planning in U-curve environment. (a) Real environment (b) Initial pose (c) Set the target pose (d) Planned path.",
        "texts": [
            " Also the target position and posture are set in rviz (Figure 17(c)), and the collision-free path generated by the algorithm is shown in Figure 17(d). Right angle bend and narrow passage environment In the indoor environment, the wooden boards are used to build the right-angle bends, and the obstacles are placed at the corners to reduce the width of the passage, which simulates the scene of passing through the door, as shown in Figure 18(a). The environment map is established (Figure 18(b)), the target position and posture are set (Figure 18(c)), and a path is planed by using IBA (Figure 18(d)). The U-shaped curve environment is built as Figure 19 (a) and the environment map is also established. The target position and posture are set, and the path is planed by using the proposed global path planning algorithm, shown as Figure 19(d). Through the above four experiments in different environment, it can be seen that in the barrier-free environment, the global path approximates a straight line. For the static obstacle environment, the improved algorithm in this paper can avoid obstacles and plan a smooth path. For the complex scenes, the results of experiments (3) and (4) show that the global path planning algorithm based on IBA can plan a reasonable path. The path can avoid obstacles and is relatively smooth, which verifies the feasibility and effectiveness of the improved algorithm in this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000774_cjme.2013.03.573-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000774_cjme.2013.03.573-Figure5-1.png",
        "caption": "Fig. 5. Machine-tool settings in pinion generation",
        "texts": [
            " Theoretically, the motion curve can be formulated in Taylor series up to the second-order as follows: 2 2 20 1 1 10 2 1 10( ) ( ) ,a a\u03c6 \u03c6 \u03c6 \u03c6 \u03c6 \u03c6     (3) where 10\u03c6 and 20\u03c6 are the initial rotational angles of the pinion and gear, respectively. a1 is the ratio of the pinion tooth number N1 to the gear tooth number N2 at the mean contact point. a22 is the second derivative of the motion curve. The above second-order motion curve is realized by correcting the HOC in modified roll motion[23], which can be formulated as 2 3 4 4 1 1c c1 c1 c1 c1 c1( ),m C D E F\u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8     (4) where \u03c81 and \u03c8c1 are the rotational angles of the pinion and the cradle as in Fig. 5(a); m1c is the ratio of cutting; , , , C D E F are the HOC. Taking the RTCA motion curve as the target, the initial machine-tool settings and the HOC are formed. Since the pitch angle of the concerned gear in this paper is less than 70\u00b0, the pinion is usually generated with modified roll motion according to Gleason Works practice. Six coordinate systems are used in pinion tooth surface generation as shown in Fig. 5. CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7577\u00b7 Coordinate system Sm1 is rigidly attached to machine frame. Movable coordinate system Sc1 is used to represent the angular position of the cradle. Coordinate system Sp is used as a reference to describe the angular position of the head cutter. Coordinate system S1 is rigidly connected to the pinion. Auxiliary coordinate systems Sa1 and Sb1 are used to help the installation of the pinion with respect to the cradle and the head cutter. As can be seen from Fig. 5, there are five parameters in the machine-tool settings for the generation of the pinion tooth surfaces. They are the installment angle q1, the radial setting Sr1, the vertical offset Em1, the sliding base XB1, and the machine center to back XD1. They are determined by the local synthesis according to the meshing performances at the mean contact point. When the head cutter revolves about its own axis, the straight blade formed a cone as in Fig. 6, which acts as the tooth surface of the cutter. The equations for the cone surface and its unit normal in the coordinate system Sp are represented as follows: p p p p p p p p p p p p p ( sin )cos ( , ) ( sin )sin , cos R u u R u u \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1             r (5) p p p p p p p cos cos ( ) cos sin , sin \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1             n (6) where up and \u03b8p are the tooth surface parameters; \u03b1p is the pressure angle; Rp is the point radius of the head cutter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003256_1.c035739-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003256_1.c035739-Figure9-1.png",
        "caption": "Fig. 9 Schematic for axis deviation.",
        "texts": [],
        "surrounding_texts": [
            "As shown in Fig. 8, to simulate the axis deviation of the landinggear retractionmechanismcaused bywing deformation and assembly error, the position of up-node rotation axis on the landing gear is adjusted during the test. The specific implementation procedure of axis deviation is as follows: four sets of wedge-shaped plates (as shown in Figs. 9 and 10) are arranged, where each of the large and small plates is matched into one set. The small wedge-shaped plates are weld connected with the up-node fixture, whereas the large wedge-shaped plates are attached to the bevels of small wedgeshaped plates and locked by bolts. When artificial generation of axis deviation is necessary, the bolts are loosened and the large wedgeshaped plates are knocked from two sides. After the desired deviation is achieved, the bolts are tightened again. The test content is fundamentally consistent with the hinge clearance test."
        ]
    },
    {
        "image_filename": "designv11_30_0003522_s00542-020-05045-8-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003522_s00542-020-05045-8-Figure1-1.png",
        "caption": "Fig. 1 Mechanical design of 3D multiscale gripper. a Single-piece compliant mechanism constructed from single-piece compliant 5-beam mechanism. b 3D model of a Multiscale grasping device realized from building block beam mechanism",
        "texts": [
            " The scalability study is carried out to understand how gripper structural response changes when the size of the gripper changes. Finally, Sect. 4 introduces an extensive experiment to study a rapidly prototyped gripper under electromagnetic actuation. These results are preliminary and further investigation are undergoing to understand the tradeoffs between homogeneity and linearity of gripper. A state-of-the-art kinematic for a grasping device is developed from a compliant 5-beam mechanism which is used as a building block to construct a multi-finger, handlike manipulator with the example shown in Fig. 1. In general, a gripper with a higher number of fingers can be configured by arranging a multiplicity of n-number of 5-beam mechanisms into a circular pattern creating a three dimensional EOAT for pick-n-place manipulation of irregular size objects. Simultaneous or independent linear actuators that bend fingers produce the open and close movements of the gripper fingertips. One aim is to develop a configuration for the manipulation of meso (millimeter) to microscale objects with a submicron displacement resolution. The meso scale manipulation is obtained by a linearly actuating node (n8)\u2014labeled in Fig. 1a\u2014in an upward and downward direction. This produces coarse displacements suitable to handle millimeter size objects with a large magnitude contact force that is generated between the fingertip and object being grasped. One possible actuation source to allow coarse manipulation is for an elastic structure to be driven by a bidirectional EM actuator. In this mode, the fingertips of the grasping device would open and close simultaneously and in a symmetrical movement, providing that first, all mechanisms are arranged in a symmetrical patter, and second, beams L2 are all anchored to one node (n8)",
            " 2b could be rapidly prototyped at once from elastic material and then packaged with two sets of actuators: uni or bidirectional actuator (FPE)\u2014which correspond to vector (9)\u2014and one uni or bidirectional actuator attached to a moving shuttle (D), which also corresponds to vector (10). Both are used to control the contact force and tip displacement at the tip of beam (n11). The elastic beams could be embedded with a surface gauge sensor at a stressed region where the measurement can be used in realtime feedback control of the contact force and displacement at the fingertips. Such sensors may include distributed strain gauges or thin film deposited on the surface of the flexure beams. The compliant beam (L5) in Fig. 1a provides a means to suspend the moving shuttle in a vertical direction, and to suppress lateral tilts. It also prevents the contact friction between the moving parts and help self-align and selfstabilize the fingertips during operation. Moreover, it could be used as a design parameter for matching the overall compliances between the gripper and shuttle actuators. Therefore, the static and dynamic characteristic performance of the gripper can be tuned to meet manufacturing and design specifications under given actuation sources",
            " Finally, there are mechanical limiting stops inside the structure that define a safe operating range for the coarse deflection during the pull-down or push-up modes. The contact force between fingertips and the object being grasped can be regulated via adjustment of gripper compliance or by coating the fingertips (n12). The proposed device is designed for manufacturing with a short cycle time due to its continuum structure that allows it be rapidly printed in a single job from desired elastic material. This paper will study the design model in Fig. 1 with PE and electromagnetic (EM) actuation sources for general macrosize and micro-size manipulation. The design parameters that describe the performance characteristic of a gripper are contact force magnitude, bandwidth, range and stroke (Buzuayene 2008). The contact gripping force is the force applied by the gripper on the object and transmitted from actuators to the fingertips through beam mechanism. This contact force should be sufficient enough to hold the object from slipping especially during dynamic motion of gripper",
            " This is obtained by studying the maximum tip displacement corresponding to closing and opening. The third objective is to study how the gripper performance changes as the geometry is scaled. Finite element static-structural analysis is performed on a 1-D line model constructed from 2-node structural beam elements with circular cross sections. The line model provides accurate results as compared to a 3D solid model (Boeraeve 2010). The material properties of the structure and the geometry of the 1-D model in Fig. 1a are given in Table 2. The angle between L2 and L3 is a \u00bc 180 . Piezoelectric and EM forces are applied at nodes (7) and (8), respectively. The 1-D line model is created in ANSYS and the material properties are assigned. The model supports in Fig. 1a have no displacements or rotation. The system structural response is assumed to be in the linear elastic region, and therefore a superposition model applied to study where the total geometrical advantage GA = GA1- ? GA2 results from the two input displacements. The first analysis studies the coarse displacement under open and closed mode where node (7) is fixed, and then a bidirectional input displacement is applied at node (8) with a magnitude of Ui1 = 3 mm. This corresponds to a force of 0.04 N for a typical EM solenoid actuator",
            " In ANSYS Workbench a and Li are entered as driving input parameters, and r is entered as output parameter. The objective functions Uo1 and Uo2 are maximized using Multi-Objective Genetic Algorithm (MOGA) Optimization method in ANSYS Workbench. This method is used because the system is linear, therefore Eq. (1) is equivalent to maximizing two objective functions. The size of each link is constrained such that Li Lio, where Lio are set equal to link dimensions in Table 2. The location of the node (6) in Fig. 1a is chosen as an input parameter, where the manipulation of its x (horizontal) and y (vertical) positions produces various configurations. The capability of the gripper is measured by evaluating the maximum radius of the object that the gripper can hold. Based on initial geometry it is given by R \u00bc Xp \u00feW X; \u00f03\u00de where R is the maximum radius of the object handled by the gripper. Xp is the X-directional deformation at the chosen candidate point, and X is x-coordinate of optimized node (6). W is the width of the mechanism in Fig. 1a and it is set to 9 mm. The optimization of the objective function is computed for two sets of input parameters: The first set is selected such that node (6) is bounded to {9 x 0; 18 y 0}. The optimized coordinates of node (6) that satisfies Eqs. (1) and (2) are x \u00bc 8:1602mm and y \u00bc 3:15mm. This yields the following optimized lengths L 1 \u00bc 7:8225; L 2 \u00bc 3:35569; L 3 \u00bc 14:85 mm. The calculated values for Xp and R are 13.218 mm and 14.0578 mm, respectively. Similarly, the second set is selected such that node (6) is bounded to {18 y 0}"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000671_j.ifacol.2015.07.055-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000671_j.ifacol.2015.07.055-Figure1-1.png",
        "caption": "Figure 1. Healthy PMSM stator diagram a) and stator with short winding fault diagram b)",
        "texts": [
            " Electrical torque TE compounded from reluctance component and electromagnetic component is generated according equation: TE = PP ( 1 2 iabc T dLabc d\u03b8 iabc + iabc T eabc \u03c9 ) . (4) 3. WINDING INTER-TURN SHORT FAULT AND OPEN PHASE FAULT MODELLING Principle of inter-turn short winding model was described by Liu (2006), Gandhi et al. (2011) and Vaseghi et al. (2011) based on analogy to transformer inter-turn faults. However all of these authors use faulty models only for non-saliency PMSM type, which has no inductance fluctuation to rotor angle (Ld = Lq). shown in figure 1 b). In healthy state, this couple of windings is connected in series. On the other hand, in fault state, the fault part of winding is short circuited by switch. Ratio of short circuited turns to total number of turns per phase is marked as \u03c3, that denotes severity of inter-turn fault. Faulty phase have to be included into healthy model, that was mentioned in introduction. So vectors in equation (1) have to be extended by uf , if and ef . Equation (1) can be rewritten with extension of faulty phase f as uabcf = Rabcf iabcf + dLabcf iabcf dt + eabcf "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure6.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure6.2-1.png",
        "caption": "Fig. 6.2 Linear machines with dual Halbach array",
        "texts": [
            " . . . . . . 93 Figure 5.21 Field variation at the center of external magnet area . . . . . . . 94 Figure 5.22 Magnetic field variation at z \u00bc 0 mm . . . . . . . . . . . . . . . . . 95 Figure 5.23 Flux density of air area at internal and external air space . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 96 Figure 5.24 Force variation of four back iron patterns . . . . . . . . . . . . . . 96 Figure 6.1 Experimental investigation on tubular linear machine . . . . . . 100 Figure 6.2 Linear machines with dual Halbach array . . . . . . . . . . . . . . 100 Figure 6.3 Structure of the research prototype . . . . . . . . . . . . . . . . . . . 101 Figure 6.4 Experimental testbed for magnetic field measurement . . . . . . 102 Figure 6.5 Magnetic field variation versus r at z \u00bc 0 mm. . . . . . . . . . . 103 Figure 6.6 Magnetic field variation versus r at z \u00bc 9 mm. . . . . . . . . . . 104 Figure 6.7 Magnetic field variation versus z at r \u00bc 10:5 mm . . . . . . . . 105 Figure 6.8 Magnetic field variation versus z at r \u00bc 11:5 mm ",
            " This chapter begins with description of research prototype, magnetic field measurement procedure of the tubular linear machine, and the corresponding data processing and analysis. The force and the armature reaction measurements are carried out on the machine and compared with the theoretical models. The experimental results confirm the proposed analytical model for further control study of the linear machine. A tubular linear machine with dual Halbach array is developed for experimental investigation as shown in Fig. 6.2. The windings are mounted on the mover that in turn is fixed on the two guiders. The guiders can slide back and forth on the linear bearing. The bearings are installed on the covers of the stator on which the PM array is mounted. The structure of the research prototype is shown in Fig. 6.3 and design parameters are given in Table 6.1. The maximum linear stroke is 36 mm. PMs in the machine are sintered NdFeB35 with Brem = 1.2T and \u03bcr = 1.0997. The radially magnetized PMs are replaced with segments of diametrically magnetized sector PMs for the convenience of manufacturing and cost reduction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000046_ukrcon.2019.8879957-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000046_ukrcon.2019.8879957-Figure6-1.png",
        "caption": "Fig. 6. Visualization of process of manipulation in the program RobotStudio environment with the use of GS on the pneumatic grippers",
        "texts": [
            " 3 a,c) coordinates of the center of masses gripping systems will have values (0, 0, 130) and the mass of m=5 kg, and for a vacuum gripping system (Fig. 3 b) (0, 0, 131) and m=4.56 kg. For transportation with optimization of orientation with Bernoulli gripping devices and vortex (Fig. 3 d,f) coordinates of the center of masses gripping will have systems values (127, 0, 206) and mass of m=5.2 kg, and for a vacuum gripping system (Fig. 3 e) (128, 0, 206) and m=4.7 kg. Let's carry out modeling of process of transportation in the program RobotStudio environment for all parameters of gripping systems (Fig. 6,7). Apparently from the schedule (Fig. 7), during performance of transport operation with orientation optimization by the industrial robot more energy, than without orientation optimization is spent for 43%. It is connected with the fact that for change of orientation of a final effector engines of a brush of the industrial robot which are not used during transportation without orientation optimization are used. Now we will consider total power costs of the industrial robot and gripping systems on transportation of a subject to manipulation (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000728_s1023193514060081-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000728_s1023193514060081-Figure1-1.png",
        "caption": "Fig. 1. Schematic illustration of sensing element in the car bon paste matrix (on the right) and CPE as an indicator electrode (IE). General protocol to determine aluminum ion (on the left). The sensor is placed into the electro chemical cell and the free aluminum ion is quantitatively measured by potentiometric method.",
        "texts": [
            " After that, paraffin is melted in a dish by a cal ibrated heater at 45\u201350\u00b0C. Subsequently the mixture of graphite and OEP is added to the melted paraffin and homogenized with a stainless steel spatula. This prepared paste is used to fill electrode which made of a glass tube (ID of 5 mm and a height of 12 cm). A cop per wire (15 cm length, 0.4 mm outer diameter) is inserted to the composite at the end of the electrode body. After cooling at room temperature, the electrode surface is burnished on a paper to remove the residual particles (see, Fig. 1). Afterward, the CPEs are condi tioned in the 10\u20133 M aluminum ion solution for 24 h at pH 5. 556 RUSSIAN JOURNAL OF ELECTROCHEMISTRY Vol. 50 No. 6 2014 MAJID SOLEIMANI, MAJID GHAHRAMAN AFSHAR All electromotive forces (EMF) are measured at room temperature (21 \u00b1 5\u00b0\u0421) in stirred solutions. A double junction Ag/AgCl/3 M KCl/1 M LiOAc refer ence electrode is used as a reference electrode. Activity coefficients are calculated using the Debye Huckel approximation and the potential measurements are corrected by using the Henderson equation to deter mine the liquid junction potential. The experimental pH values are recorded in parallel with a calibrated glass pH electrode (see, Fig. 1). About 100 mL of tap water sample is adjusted to an optimum pH 5 by adding diluted nitric acid. The CPEs are inserted in to the real water sample to deter mine aluminum ion concentration and subsequently aluminum solution is added at 4 \u00d7 10\u20136 M level of con centrations. We have achieved substantially improved sensor selectivity with incorporating OEP to the composition of CPE. Selectivity may be understood with the metal ligand reaction between aluminum ion and OEP. The new potentiometric aluminum sensor consists of OEP as an ionophore based on CPE"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003147_j.engstruct.2020.110892-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003147_j.engstruct.2020.110892-Figure15-1.png",
        "caption": "Fig. 15. Vertical loading assembly.",
        "texts": [
            " The top plate of the frame is attached to a guideway at each corner as shown in Fig. 13 to allow vertical movement of the end plate of the IMT connected to it under the axial compression. Four weights4 hung by nylon cords are used to balance the effect of weight of the top plate as shown in Fig. 4. The bottom plate is connected by bolts to the columns of the experimental frame and its vertical location can be adjusted to suit different lengths of tubes as shown in Fig. 14. The vertical force was applied by placing weights4 on the top plate as shown in Fig. 15. The transverse loads were applied by weights4 through two nylon cords connected to points at one-third heights of the tube as shown in Fig. 16. The vertical locations of the transverse loads can be adjusted for the tubes of different lengths by moving the slide blocks connecting the lockable guideways and the nylon cords as shown in Fig. 16. The vertical displacement at the top end and transverse displacement at mid-height of the tube were measured with Panasonic HGC1100 laser sensors (with an accuracy of 70 \u03bcm) as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000898_icca.2013.6564877-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000898_icca.2013.6564877-Figure1-1.png",
        "caption": "Fig. 1 Improved de-icing tool",
        "texts": [
            " The robot may sometimes be blocked in two cases: a) Ice blocks which are cut into may be big enough, when the ice is very hard, to get stuck between the cutting tool and the front wheel; b) ice could not be cleared all away after rushes with de-icing tool. Remained ice may block on the edge of the hole in the middle of the cutting tool, and prevent the robot from going backward ready for next rush. Many subtle changes are made Extended Applications of LineROVer Technology Hangzhou, China, June 12-14, 2013 978-1-4673-4708-2/13/$31.00 \u00a92013 IEEE 1415 on the de-icing tools, as shown in Fig. 1, to solve the problems. The clearance between the cutting tool and the front wheel is increased to ensure the broken ices could drop from the interspaces. And the hole in the middle of the cutting tool is reduced. The forward edge of the hole is made angular. It is sharp enough to cut up ices when the robot rushes forward. However, the other side of the hole is smoothed in the edge. It is guaranteed that when the robot goes backward, it could slide along the smooth edge out of remained ices. Transmission lines are iced up usually in foggy or frozen rainy weather"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003606_01691864.2020.1835532-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003606_01691864.2020.1835532-Figure1-1.png",
        "caption": "Figure 1. Basic ideas for CoM estimation. (a)CoM kinematics. (b)Linear momentum. (c)Angular momentum (c-1: direct use of angular momentum, c-2: ZMP-based method.)",
        "texts": [
            " This is because the humanoid robot is more strongly expected to apply to the multi-contact situation than before. This paper first summarizes three basic ideas for CoM estimation and refers to those disadvantages. In this section, since both the CoM estimators for humanoid robots and humans are based on common ideas in many studies, we also cite the CoM estimators proposed in the field of anthropometry. Most common and widely-used method to estimate the CoM is the kinematics based on the inertial parameter of each link as shown in Figure 1(a). Let pG = [xG yG zG]T \u2208 R 3 be the CoM with respect to the inertial frame , it is represented as a weighted sum of the CoM of each link, namely, pG = 1 m Nl\u2211 i=0 mipG,i, (1) m = Nl\u2211 i=0 mi, (2) where m is the total mass of the robot, Nl is the total number of links excluding the base-link, and mi is the mass of ith link. pG,i \u2208 R 3 is the CoM position of ith link with respect to . Hereafter, the subscript i = 0 means the base-link frame of the robot 0. Furthermore, pG,i is expressed as pG,i = p0 + R0 0pG,i, (3) 0pG,i =0 pi + 0Ri ipG,i, (4) where p0 \u2208 R3 and R0 \u2208 SO(3) denote the position and attitude of 0 with respect to , respectively",
            " In many cases, the designed link geometric parameters are available, and joint displacements are able to be measured by the joint encoder accurately. On the other hand, it is necessary to estimate the base kinematics by only sensors attached to the robot. This issue will be described in Section 4. Finally, because of errors between the inertial parameters of the CAD model and of the real robot, those identification is required (e.g. [32,38,39]). Another method to estimate the CoM focuses on the linear momentum as shown in Figure 1(b). The linear momentum l \u2208 R 3 is represented as l = mp\u0307G. (8) Its time-derivative corresponds to the sum of total force acting on the robot f \u2208 R 3 and the gravity, namely, l\u0307 = mp\u0308G = f \u2212 mg, (9) where g = [0 0 g]T \u2208 R 3, and g is the acceleration due to the gravity. Since f is the sum of the force acting on all contact point, it is represented as f = NC\u2211 j=1 f j, (10) where NC is the total number of contact points, and f j = [fxj fyj fzj]T \u2208 R 3 denotes the force acting on the jth contact point with respect to ",
            " Even if one part is lighter and smaller than the other as in the skin sensor [40], this issue remains. Additionally, because most skin sensors measure normal forces (pressures), the external wrench reconstructed from those sensors can contain errors due to the lack of the measurement of shear stress. Therefore, the integration error is accumulated in (12) and (13). Finally, the estimation is necessary to know pG0 and vG0, and thus those estimation errors remain. Third method focuses on the angular momentum as shown in Figure 1(c-1). The angular momentum k \u2208 R 3 is represented as k = Nl\u2211 i ( Ri iIiRT i \u03c9i + mi(pG,i \u2212 pG) \u00d7 p\u0307G,i ) , (14) Ri = R0 0Ri, (15) \u03c9i = \u03c90 + R0 0\u03c9i, (16) where iIi \u2208 R 3\u00d73 means the inertia of ith link with respect to i. From the time-derivative of k, the moment equilibrium between all contact point and the CoM is represented as k\u0307 = nG = NC\u2211 j=1 ( (pCj \u2212 pG) \u00d7 f j + nj ) , (17) where nG = [nGx nGy nGz]T \u2208 R 3 and nj \u2208 R 3 mean the moment around the CoM and the jth contact point, respectively, and pCj = [xCj yCj zCj]T \u2208 R 3 is the position of the jth contact point with respect to . pCj and nj are written as pCj = p0 + R0 0pCj , (18) nj = R0 0RCj Cjnj, (19) where 0pCj \u2208 R 3 means the position of the jth contact point with respect to 0, and Cjnj \u2208 R 3 is the moment around the jth contact point represented in Cj . (17) is rewritten as n = pG \u00d7 f + nG (20) = NC\u2211 j=1 ( pCj \u00d7 f j + nj ) . (21) Instead of (20) and (21), as shown in Figure 1(c-2), many studies [8,10,13,19\u201321,26,29,41\u201344] utilized the relationship between the ZMP and the CoM represented as nZ = ( pG \u2212 pZ ) \u00d7 f + nG (22) = NC\u2211 j=1 (( pCj \u2212 pZ ) \u00d7 f j + nj ) , (23) where pZ = [xZ yZ zZ]T \u2208 R 3 is the ZMP position with respect to , and nZ = [0 0 nzZ]T is the moment around the ZMP. Since zZ can be determined arbitrarily, xZ and yZ are computed by the component of x and y in (23), namely, xZ = \u2211NC i=1 ( \u2212nyj + xCj fzj \u2212 (zCj \u2212 zZ)fxj ) \u2211NC i=1 fzi , (24) yZ = \u2211NC i=1 ( nxj + yCj fzj \u2212 (zCj \u2212 zZ)fyj ) \u2211NC i=1 fzj "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-13-1.png",
        "caption": "Figure 6-13. Step 6 to obtain the propulsion or allocation matrix",
        "texts": [
            " The direction in which the rest of the fingers point in a closed fist is the positive direction of rotation). Just as the previous step influences vehicle\u2019s movement and control, this will also depend on the lever test of the remote control. Therefore, if the levers are inverted with respect to a desired movement, the proposed direction of rotation must be modified. Notice that the proposed situation changes for the design of omnidirectional vehicles, and it is convenient that all the rotating axes are selected according to the rule of the right hand. See Figure\u00a06-13. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 256 Step 7: Answer the following questions. 1. Which motors provide direct displacement on the X axis of the drone? In this case, none, Fx\u00a0=\u00a00, because by being well balanced, each propeller only transmits a vertical force and a torsional moment. You can also assume that the vehicle is well balanced, which means that the whole mass of the vehicle is concentrated at its geometric center where the autopilot reference mark is placed and its arms have the same length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003663_012104-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003663_012104-Figure8-1.png",
        "caption": "Figure 8. The 3D-stress analysis results.",
        "texts": [
            " Exact value of resultant stresses did not play a significant role in the simulation results. Mechanical characteristics of materials were: of steel - mass density 7850 kg/m3, Young\u2019s modulus 2.1\u00b7105 MPa, Poissons ratio 0.3; of bronze - mass density 8960 kg/m3, Young\u2019s modulus 1.1\u00b7105 MPa, Poissons ratio 0.35. Parameters of mesh: 9\u00b710-5 m approximate global size of seeds, 0.1 maximum deviation factor, linear geometric order C3D8R element type (an 8-node linear brick). The results of numerical simulation (figure 8) proved that mechanical stresses in the specimen type \u21161 had their ultimate values in the interface between steel and bronze (the red circles in the figure 8). The model in figure 8 is demonstrated with local angular cross section which shows distribution of stresses inside the cylinder. Blue area in the base of cylinder is a zone with the lowest value of mechanical stresses (what was certainly predictable because of boundary conditions). A light-green area in the figure 8 demonstrates a low-stressed zone. RusMetalCon 2020 IOP Conf. Series: Materials Science and Engineering 969 (2020) 012104 IOP Publishing doi:10.1088/1757-899X/969/1/012104 A difference between thermal expansion coefficients of the materials along with rapid intermetallic growth, increase of microhardness (up to 223 HV), increase of elasticity modulus (up to 36.7 GPa), stress concentration and local embrittlement of the FG laser deposited specimens with two sharp transitions created of stainless steel AISI 316L and aluminium bronze in the DMT mode via the straight joining scheme lead to cracking on a border between stainless steel and aluminium bronze"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003618_ever48776.2020.9243048-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003618_ever48776.2020.9243048-Figure1-1.png",
        "caption": "Fig. 1. Structure of the investigated PMSM.",
        "texts": [
            " At steady state, although the rotor is synchronized with the fundamental component of the stator magnetic motive force (MMF), time and space harmonics of the MMF which are induced by the armature currents and the unevenness of the air-gap permeance can still cause certain rotor eddy current, and consequent power loss. Since the magnetic flux density is the product of the MMF and the permeance, the magnetic field fluctuation in the rotor is the result of either the MMF (produced by the magnets as well as the armature current) variation, or the magnetic circuit permeance variation. To separate the rotor eddy current loss caused by diverse reasons, the investigated PMSM with ratings of 80 krpm and 1.2 Nm, shown in Fig. 1, has been run under no load (i.e., no armature current), rated load with ideal sinusoidal current, and rated load with PWM applied voltage source inverter Authorized licensed use limited to: Carleton University. Downloaded on June 02,2021 at 06:00:22 UTC from IEEE Xplore. Restrictions apply. (VSI) conditions, separately. The motor has 3 phases, 2 poles and 12 slots, and each phase has 2 short-pitch overlapping coils. Clearly, even the pure sinusoidal current will make space harmonics in the armature MMF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001537_roman.2015.7333585-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001537_roman.2015.7333585-Figure1-1.png",
        "caption": "Fig. 1. Tension-sensitive elastic thread: (a) An electro-conductive yarn is a mixture of conductive (30 %) and non-conductive (70 %) filaments. Extension of the yarn increases the density of the conductive filaments, reducing the resistance of the yarn. (b) A fiber sensor consists of a nonconductive core and two electro-conductive yarns winding around the core in the opposite directions.",
        "texts": [
            " For example, a silicone tube filled with eutectic gallium indium was used to measure the bending of robotic fabrics [17]. In this paper, we propose a fiber sensor with coupling structure attachable to the membrane of a McKibben actuator to measure actuator shrinkage. We have applied electro-conductive yarns to a fiber sensor. This electro-conductive yarn consisted of a mixture of conductive stainless filaments (30 %) and non-conductive polyester filaments (70 %). Extension of this yarn increased the density of conductive filaments (Fig. 1(a)), reducing the resistance of the yarn. Unfortunately, the resistance saturated when the extensional strain exceeded 1 %, which is insufficient for measuring the shrinkage of a McKibben actuator. We therefore utilized coupling structure of electroconductive yarns for fiber sensors. This structure consists of one non-conductive polyurethane core and two electroconductive yarns winding around the core in the opposite directions to each other (Fig. 1(b)). Extension along the yarn is smaller than that along the core, suggesting that the extension of this fiber sensor can be measured in a wider range. We found that resistance of this fiber sensor decreased monotonically as its extensional strain increased up to 20 %, suggesting that this sensor can measure extension of up to 20 % of its length. If a fiber sensor is attached onto the membrane of a McKibben actuator (Fig. 2), with both ends of the sensor fixed on the membrane, then the distance between the two ends along the sensor can be determined by measuring the resistance between the ends"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002808_0959651819899267-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002808_0959651819899267-Figure2-1.png",
        "caption": "Figure 2. DC motor circuit diagram.",
        "texts": [
            " The main and tail DC motors are identical with different mechanical loads. Both the DC motor rotors are attached to the corresponding propellers whereas the stators (housings) are attached to the support beam ends, as shown in Figure 1. The active torques form the motors are applied to the propellers and the reactive torques act on the support beam through stators. These reactive torques cause cross-coupling between the yaw and pitch. The circuit diagram of the permanent magnet DC motor is shown in Figure 2 where parameters with subscripts m and t correspond to the main and the tail motors, respectively. The differential equations governing the electrical and mechanical domains of the permanent magnet DC motor are, respectively, given as Lam=t diam=t dt =Vm=t Kam=t vm=t Ram=t iam=t \u00f01\u00de Jm=t dvm=t dt =Kam=t iam=t TLm=t Cvm=t vm=t \u00f02\u00de where Vm/t is the input terminal voltage, Lam/t and Ram/t are the motor armature inductance and resistance, respectively, iam/t is the armature current, vm/t is the angular velocity of the rotor, Kam/t is the torque constant of the motor, Jm/t is the mass moment of inertia of the rotor about the axis of rotation, Cvm/t is the viscous damping, and TLm/t is the resisting torque due to the drag force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001233_amr.874.57-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001233_amr.874.57-Figure2-1.png",
        "caption": "Fig. 2. The 3D kinematic model (a) and kinematic structure (b) of the 3-RRPRR hydraulic TPM: R \u2013 revolute joints, P \u2013 prismatic joints, TCP \u2013 tool center point.",
        "texts": [
            " The hydraulic cylinders was placed on the fixed base about circumradius R = 250 [mm], the rod cylinders are connected by the joints with the moving platform about circumradius r =130 [mm]. The maximum length of the hydraulic axis carries out Limax = Limin + h, where Limin = 425 [mm] is a initial length. The closed-loop kinematic chains of the hydraulic TPM create structure 3-RRPRR, in which revolute joints R and prismatic joints P step out. The 3D kinematic model and kinematic structure of the 3-RRPRR hydraulic TPM was shown in Figure 2. Kinematic of the hydraulic parallel manipulator The computational kinematic model of 3-axis hydraulic parallel manipulator shown in Fig.3. For the contour error of the 3-axis hydraulic parallel manipulator the 6-DoF kinematic model have been considered. It includes both the platform position variables xp, yp, zp and also rotational RPY angles , ,  . The RPY (Roll-Pitch-Yaw) angles defined with respect to three successive rotations about the fixed X, Y, Z axes [4]. The problem of inverse kinematics is to find the hydraulic cylinders elongation Li, given the position (xp, yp, zp) and orientation error (, , ) of the TCP"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003147_j.engstruct.2020.110892-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003147_j.engstruct.2020.110892-Figure12-1.png",
        "caption": "Fig. 12. Structure of the fixed end of the IMT.",
        "texts": [],
        "surrounding_texts": [
            "The detailed structures of the simply-supported and fixed ends of the IMT in the experiment are designed as presented by Figs. 11 and 12. The top plate of the frame is attached to a guideway at each corner as shown in Fig. 13 to allow vertical movement of the end plate of the IMT connected to it under the axial compression. Four weights4 hung by nylon cords are used to balance the effect of weight of the top plate as shown in Fig. 4. The bottom plate is connected by bolts to the columns of the experimental frame and its vertical location can be adjusted to suit different lengths of tubes as shown in Fig. 14."
        ]
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure1.9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure1.9-1.png",
        "caption": "Fig. 1.9 Structure of the moving-coil-type linear DC motor [23]",
        "texts": [
            "4 Linear machines for applications of Industrial Automation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 Figure 1.5 Linear machines for applications of rail transportation . . . . . 6 Figure 1.6 Winding arrangements for tubular PM linear machines [19] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 Figure 1.7 An axial flux permanent magnet motor [21] . . . . . . . . . . . . 9 Figure 1.8 Structure of the moving-coil-type linear DC motor [22] . . . . 10 Figure 1.9 Structure of the moving-coil-type linear DC motor [23] . . . . 10 Figure 1.10 Schematic of reciprocating linear generator system [24] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.11 A linear vibration-driven electromagnetic micro-power generator [26] . . . . . . . . . . . . . . . . . . . . . . . . 11 Figure 1.12 An improved axially magnetized tubular PM machine topology [28] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 Figure 1",
            " The structure is illustrated in Fig. 1.8. It consists of a coil, permanent magnets, yokes, a coil bobbin held by the arms of a coil holder and linear bearings. There are 16 pieces of permanent magnets, and they are fitted along the inner surface of the outer yoke and magnetized radially. The yokes are solid steel blocks. The inner yokes are separated into four pieces to provide space for the arms of the coil holder. Kim et al. presents a tubular linear brushless permanent magnet motor as shown in Fig. 1.9 [23]. It has a slotless stator to provide smooth translation without cogging. In this design, the magnets in the moving part are oriented in an NS-NS SN-SN fashion which leads to higher magnetic force near the like-pole region. Wang et al. describes the design and experimental characterisation of a reciprocating linear permanent magnet generator developed for on-board generation of electrical power fix telemetry vibration monitoring systems as shown in Fig. 1.10 [24, 25]. The two axially magnetised sintered NdFeB magnets and the mild-steel pole pieces give rise to an essentially radial magnetic field in the region occupied by the generator winding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001392_s11721-013-0084-9-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001392_s11721-013-0084-9-Figure2-1.png",
        "caption": "Fig. 2 Change of orientation due to interaction with neighbors. The decision making individual is located at x and is moving in direction \u03c6. (a) Attraction. The reference individual at x is more likely to turn due to its interaction with s\u2032 than with s, as |\u03c6 \u2212 \u03c8 \u2032| > |\u03c6 \u2212 \u03c8 |. (b) Alignment. The individual at x is more likely to turn due to its interaction with s than with s\u2032 , as |\u03c6 \u2212 \u03b8 | > |\u03c6 \u2212 \u03b8 \u2032|",
        "texts": [
            " The Aj \u2019s are normalizing constants that make the spatial integral of each kernel equal to 1. These constants are given by Aj = \u03c0mj ( mje \u2212d2 j /m2 j + \u221a \u03c0dj ( 1 + erf(dj /mj ) )) . (4) Figure 1(b) shows the attraction kernel Kd a . Note that the three interaction zones may overlap. (ii) Orientation kernels. We discuss first the attractive orientation kernel. Suppose a decision making individual located at x is heading in direction \u03c6 and senses a neighbor located at s, within its attraction zone\u2014see Fig. 2(a). The relative location s \u2212 x makes an angle \u03c8 with the positive x axis. Due to attraction, the reference individual makes a decision to turn in order to approach its neighbor at s. The larger the difference |\u03c6 \u2212 \u03c8 |, the higher the likelihood of the decision making individual to turn. These considerations are included in the following attraction kernel1 Ko a : Ko a ( s; x,\u03c6) = 1 2\u03c0 (\u2212 cos(\u03c6 \u2212 \u03c8) + 1 ) . (5) Note that Ko a is smallest when \u03c6 and \u03c8 are the same ( x is already moving towards s, hence no need to turn), and largest when \u03c6 and \u03c8 are an angle of \u03c0 away from each other ( s is directly behind x). Referring again to Fig. 2(a), x is more likely to turn because of the attractive influences of s \u2032 compared to s. The 1 2\u03c0 factor renormalizes the kernel so it integrates to 1. Very similar considerations can be made regarding the change of orientation due to repulsion. The repulsion kernel, Ko r has the expression: Ko r ( s; x,\u03c6) = 1 2\u03c0 ( cos(\u03c6 \u2212 \u03c8) + 1 ) . (6) Compared to (5), (6) has change of sign in front of the cosine term, which reverses the likelihood of turning. 1Superscripts d and o in the interaction kernels refer to distance and orientation (angle), respectively. Subscripts r,al, a stand for repulsion, alignment and attraction, respectively. Finally, regarding the alignment orientation kernel, we refer to Fig. 2(b). The decision making individual is located at x and is moving in direction \u03c6. A neighbor s located in its alignment range moves in direction \u03b8 . The larger the relative orientation |\u03c6 \u2212 \u03b8 |, the higher the likelihood of the reference individual to turn to align with its neighbor. Similar to previous considerations, this is encoded in the following function: Ko al(\u03b8,\u03c6) = 1 2\u03c0 (\u2212 cos(\u03c6 \u2212 \u03b8) + 1 ) . (7) The six kernels from (3), (5), (6) and (7) are the building blocks for both turning rate functions \u03bb and T ",
            " Note that in the expression for Tal, the argument of wal includes the direction \u03b8 . This is because the alignment contribution to turning does not depend on the location of the neighbors, \u03c8 , but rather on their direction, \u03b8 . 2.2 Model with a field of vision In order to increase the biological realism of the model (1) from Fetecau (2011), we introduce a field of vision/blind zone of the individuals. The model presented in Sect. 2.1 suggests that, in terms of attractive interactions, the neighbors behind a reference individual (see Fig. 2(a)) have the strongest influence on it. However, most animals cannot see behind themselves and are not susceptible to sudden changes occurring behind them, especially when animals depend primarily on sight. We introduce here a blind zone that is in a consistent format to the model presented in Sect. 2.1. The field of vision/blind zone is modeled via an additional truncation kernel. We refer again to Fig. 2(a). To correctly represent the field of vision of the reference individual x, the neighbors located in front (with \u03c8 within a certain interval centered at \u03c6) should be given higher weights than the individuals located behind (|\u03c8 \u2212 \u03c6| close to \u03c0 ). We consider the following function that captures this idea: Kbz(\u03c6 \u2212 \u03c8) = 1 B ( 1 2 tanh { a [ cos(\u03c6 \u2212 \u03c8) + ( 1 \u2212 b \u03c0 )]} + 1 2 ) , (16) where a and b determine the steepness and the width of the field of vision, respectively, and B is a constant that normalizes the kernel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000943_2669047.2669054-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000943_2669047.2669054-Figure8-1.png",
        "caption": "Figure 8: Manufactured stylus.",
        "texts": [
            " Typical tablet PCs have a camera in front display, and the proposed system detects the user\u2019s view point using the camera. To give a user sensation as if the stylus enter into display naturally, the system measures the angle formed by the stylus and display. To measure the angle formed by the stylus and display, LED is attached to the stylus, and fish-eye lens is attached to the camera on tablet PC. The system estimates the angle formed by the display and stylus by capturing LED by build-in camera. A prototype of the proposed system was implemented. Figure 7 shows an overview of the prototype, and figure 8 shows the manufactured stylus. Nr was measured by the pressure sensor with the micro controller. Built-in camera was used for the viewpoint detection, and the virtual space is displayed according to the viewpoint. Table 1 shows the specification of the hardwares used in the prototype. Figure 9 shows a usage example of the system. The prototype has shown that the user feels a reaction force and a tactile sensation from the virtual object by using the system. In this paper, we have proposed a system for touching a virtual object in a tablet PC with a retractable stylus and a display"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003117_s00542-020-04893-8-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003117_s00542-020-04893-8-Figure6-1.png",
        "caption": "Fig. 6 Proposed yoke design. Left yoke represents the case when kt \u00bc 180; n \u00bc 4 and right one is for the case when kt \u00bc 540; n \u00bc 1. The roller tracks which are highlighted as green and blue are different from each other (color figure online)",
        "texts": [
            " 3 Scotch-yoke mechanism for calculating the roller track shapes of yokes in VSY-SEA the tracks of Fig. 5 were used for simulations and experiments. The value of kt was fixed to 180 and r was calculated by substituting 1, 2, and 4 into n. Figure 5 is shows the calculated results. To drive in both directions of the SEA, the roller track of the yoke was designed by making the graphs of r in Figs. 4 and 5 symmetric.To simplify the equation in Sect. 2, the roller was regarded as a point, and when designing the yoke, 5 mm offset was applied considering the radius of the roller. Figure 6 shows the shape of the designed yoke, and Fig. 7 shows the design shape of the completed VSYSEA. The central shaft is the output shaft and the load arm will be connected to it. The outer case of VSY-SEA acts as the input shaft and the actuator will be connected to here. When the SEA including the input shaft and the load arm rotates and the SEA receives load torque by external factors, the output shaft rotates in the opposite direction of the input shaft rotation. Then the roller which is attached to the output shaft pushes the yoke"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure7-1.png",
        "caption": "Fig. 7. Proposed 6S-13P C-Core CPSFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " Analysis reveals that conventional 6S-10P E-Core SFPMM and 6S-10P C-Core SFPMM retain same PM volume as that of 12S-10P E-Core SFPMM and increased slot area, however author fails to compensate effects of leakage flux. Until now, many researchers tried to reduced PM as much as possible and suppress flux leakage but unfortunately both effects are not considered at a time. In this paper, alternate Consequent Pole SFPMM (CPSFPMM) with partitioned PMs are introduced which further reduce PM usage and suppress flux leakage completely. For analysis and electromagnetic performance evaluation, this paper proposed three topologies of CPSFPMM (as shown in Fig. 5, Fig. 6, Fig. 7) corresponding to three topologies of conventional SFPMM as listed in table I. Based on Finite Element Analysis (FEA) proposed model has successfully reduced PM usage much more and suppress PM leakages from the end and enhance flux modulation effect. Hence proposed model reduces machine cost furthermore, retaining electromagnetic performance. The rest of the paper is organized as, section II present design parameter, design methodology and working principle of proposed CPSFPMM, section III illustrates FEA based electromagnetic performance analysis and finally, some conclusions are drawn in section IV"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002635_ecce.2016.7854747-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002635_ecce.2016.7854747-Figure6-1.png",
        "caption": "Fig. 6. Armature reaction fields extracted from the rated-load operation, under zero d-axis current control, phase current=7.5Arms.",
        "texts": [
            " Thus, the flux linkage can be flexibly changed with the aid of the current pulses fed by magnetizing coils. This merit enables the proposed machine to fulfil the specific requirement over a wide-speed-range duty-cycled operation, such as automotive applications. The influences of the armature reaction fields on the magnetization state of AlNiCo PMs are evaluated. First, the armature reaction fields are extracted from the rated-load operation with the aid of frozen permeability method [15], as shown in Fig. 6. It can be seen that the AlNiCo PMs can well resist the armature reaction fields due to the parallel magnetic circuits between PM and armature reaction fields. IV. STATOR SLOT/ROTOR POLE COMBINATIONS Based on the design theory of flux-modulation machine [10], the feasible stator/rotor pole combinations of the SCPMM are governed by ( ) = , 1, 2,3... , s s s Z mk k GCD Z Z = (14) where GCD denotes the Greatest Common Divisor, m is the phase number. On the other hand, since all the EMF vectors for the same phase are aligned as shown in (kd) for the four machines all equal to one, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002536_2.0921702jes-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002536_2.0921702jes-Figure13-1.png",
        "caption": "Figure 13. Schematic of e-/h+ generation and self-cleaning process.",
        "texts": [
            " In a typical photoelectrocatalysis process, Pt/TiO2 electrode formed a circuit with the auxiliary electrode and set the electric field direction from SCE to working electrode, so the photoinduced electrons flowed to the internal surface of the electrode while the holes of valence band shifted to the external surface. The holes then trapped electrons in the solution and gave rise to strong oxidation of reductivestate matter. In this work, the intermediates of glucose were expected to be completely mineralized to CO2 and H2O on the surface of the Pt/TiO2 electrode, as shown in Figure 13. The resistance of the Pt/TiO2 NTA electrode was substantially reduced due to the synergetic effect of the secondary calcination in N2 and loaded Pt nanoparticles, and an enhanced detection sensitivity was thus obtained in alkaline media. Meanwhile, electrocatalytic oxidation reaction occurring on the Pt/TiO2 NTA electrode toward glucose was identified to involve the transferring of two electrons and two protons. Afterwards, the contaminated electrode after longtime use was restored during photocatalysis and photoelectrocatalysis self-cleaning processes, and the electrode demonstrated a superior ) unless CC License in place (see abstract)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003029_s00170-020-05351-5-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003029_s00170-020-05351-5-Figure3-1.png",
        "caption": "Fig. 3 Geometries of the elliptic nozzle with three different minor axis lengths. a Level 1, b level 2, c level 3",
        "texts": [
            " The L9 orthogonal array of the Taguchi design has been applied in order to efficiently explore a large number of parameter changes and establish the optimal design parameters. According to our previous results [10], the elliptical shape for the blowing nozzle was selected owing to its better flow uniformity and higher momentum exchange compared to other blowing nozzle geometries. It was employed in this simulation with different aspect ratios (ARs) as the major axis was fixed to 300 mm, half of the length of the suction tunnel. Figure 3 and Table 3 show the geometries of the elliptical nozzle for three levels as well as four control factors including the widths of the blowing nozzle (Wb), the thicknesses of the suction tunnel (Ts), the suction-toplane distances (Ds), and the Reynolds numbers (Re). The simulation layout for L9 is shown in Table 4. The signal-to-noise (S/N) ratios are the log functions regarded as the prominent targets to analyze the results. The S/N ratio can be divided into three categories: the smaller the better, the larger the better, and on target (minimum variation)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure3-13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure3-13-1.png",
        "caption": "Figure 3-13. Two kinds of power distributors",
        "texts": [
            " Parameters to consider portable or external (on-board or not), maximum number of cells, light, and/or sound indicators Component keywords Lipo tester, Lipo monitor buildings or wooded areas. Chapter 3 ConCepts anD Definitions 60 Parameters to consider Multiple modules or only one, type of connector, pedestal, signal amplifier, resolution and precision, electromagnetic noise protection Component keywords eMi noise, redundant Gps, Gps accuracy, drone Gps mount Distributors are the way that more than one ESC can connect simultaneously to the main battery. They can be of an integrated circuit type or a simple current divider (a harness like an octopus connector). See Figure\u00a03-13. Chapter 3 ConCepts anD Definitions 61 Parameters to consider Maximum current supported, maximum voltage supported, number of motors to be powered, BeC, size, weight, type, electromagnetic protection Component keywords Drone power distribution board, wiring harness drone A power module is the way the autopilot is connected to the main battery; see Figure\u00a03-14. It has two outputs: one for the distributor and one for the autopilot. Always remember to verify your own battery input and distributor output"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003711_tmrb.2020.3042992-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003711_tmrb.2020.3042992-Figure3-1.png",
        "caption": "Fig. 3. Positional virtual fixture to push user\u2019s hand toward the closest point on the desired fixture curve.",
        "texts": [
            " These components were used to rotate the current orientation Ri toward Rs by ki radians per time-step. Ri = { Rs, if \u03b3 < 0.01 Ri\u22121eki\u03c9\u0302 otherwise. (11) This broadly aligned the VF with {s}, but prevented sudden changes in Rs from causing large changes in \u03c4 v f . A positional virtual fixture (PVF) was used to guide the user\u2019s hand to follow some desired path on the surface of the environment. This was achieved through a soft barrier VF with a constant stiffness ky that restricted movement transverse to the path on the MTM. A diagram of the PVF is shown in Fig. 3. The PVF was calculated with the use of a compliance frame {\u0303c}, located at the closest point, c\u0303, on the ablation path to the PSM tip. The x axis of c\u0303 was oriented along the local tangent of the ablation path, the z axis along the outward-pointing surface normal, and the y axis according to the right-hand rule. The force of the PVF was calculated in the PSM base frame, {b}, using the equation: fvf = (c\u0303Rb )\u22121 Kp c\u0303Rb(\u0303c \u2212 s\u0303) (12) s\u0303 is the point on the model environment surface that is closest to the proxy PSM. Kp is the VF stiffness gain matrix. The rotation c\u0303Rb rotates a vector in frame {b} to be represented in frame {\u0303c}. To prevent users from leaving the finite-length ablation path, a stiffness kx was applied when the user was outside the finite length of the curve. Two planes were defined in the y-z plane of the VF compliance frame {\u0303c} at either end of the ablation path, Fig. 3. A point s\u0303 was determined to belong to the curve (i.e., s\u0303 \u2208 curve) if it was between these planes. The VF stiffness matrix was defined as: Kp = diag ( kx, ky, 0 ) (13) kx = { 0, if s\u0303 \u2208 curve ky, otherwise. (14) As part of the CSA framework, a stiffness estimation module was included to semi-autonomously palpate the environment and create a stiffness map of the surface. Using the method of [35], a force-based sinusoidal palpation motion was superimposed over the user\u2019s movement commands during surface exploration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003261_0959651820935692-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003261_0959651820935692-Figure2-1.png",
        "caption": "Figure 2. Schematic of IP for modeling.",
        "texts": [
            " Acceptance of the proposed controller is assured by implementing it in real-time scenario by interacting with the LabVIEW software. To manifest the effectiveness of the described methodology, the LPV controller results are compared in real time with proportional\u2013integral\u2013derivative (PID) controller and linear\u2013quadratic regulator (LQR) controller. For the sake of identifying IP model, perusing parameters are needed as follows: Motor (holding the linked arm) position. Motor speed. Linked pendulum arm (attached to motor) position. Speed of linked pendulum arm. Figure 2 describes the schematic needed to identify the model of IP. Identifying the position of linked pendulum arm and motor position For the purpose of identification, five channels are used from NI ELVIS II. First channel is used for the input of the setup and the remaining channels are used for outputs. Motor is connected at input channel, and output channels have all sensors connected. A feedback is introduced by applying a step input to plant (Figure 3). When this feedback is introduced, system gives a step response"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-19-1.png",
        "caption": "Figure 6-19. References for obtaining the allocation matrix of a coaxial bicopter",
        "texts": [
            " This way, by combinatory theory, only for Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 261 this type of drone there are several possibilities to choose the allocation matrix and only one is correct. If your choice is wrong, the vehicle will have incorrect or even catastrophic behavior. The bicopter or PVTOL (planar vertical takeoff and landing) is a drone designed to move only along the vertical Z and horizontal Y axis. A real bicopter can move in all three dimensions but to prevent going forward and backward on the X axis, let\u2019s analyze the following example with coaxial motors, which are sold as coupled motors that share the same axis but rotate in opposite directions. See Figure\u00a06-19. w w w w 2 3 1 4 = - = - Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 262 This this way, due to the opposite direction of rotation, the thrust forces are duplicated while the torques are canceled. If you apply these criteria and analyzing as if the vehicle has a single right motor and a single left motor, you get the following: F F F x y z x y z = = = +( ) = -( ) = = 0 0 2 2 0 0 2 4 2 4 w w t w w t t Remember that the torque in X is the necessary tilt on the X axis of the drone so that it advances on the positive Y axis of the drone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003756_wre51543.2020.9307007-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003756_wre51543.2020.9307007-Figure2-1.png",
        "caption": "Fig. 2: Some of the possible locomotion configurations of the proposed robot.",
        "texts": [
            " In this way, CoppeliaSim was selected as the appropriate simulator for the confined inspection robot, given the fidelity in the robot\u2019s interactions with other objects and the ease of manipulating models and meshes directly in the virtual system [15] and facility to develop motion planning algorithms [16]. One of the EspeleoRob\u00f4\u2019s main feature is its interchangeable locomotion system, allowing adaptation to different types of terrains. The robot can drive with wheels, starshaped wheels, legs, tracks, and hybrid configurations (Figure 2). Quick-release pins allow for easy and fast exchange of the multiple locomotion systems. The robot is equipped with six sets of planetary reduction, MCD EPOS motors from Maxon Motors that have encoders and drives already integrated, 2 Bren-Tronics militarystandard batteries, Intel NUC onboard computer, Ubiquiti Rocket radio for long-distance wireless communication, and a pair of Axis cameras with high brightness LED lighting (Figure 3). Common configurations of the robot also allows for an instrumentation tower with Intel RealSense D435i and Authorized licensed use limited to: Carleton University"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003756_wre51543.2020.9307007-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003756_wre51543.2020.9307007-Figure3-1.png",
        "caption": "Fig. 3: Robotic system architecture overview.",
        "texts": [
            " The robot can drive with wheels, starshaped wheels, legs, tracks, and hybrid configurations (Figure 2). Quick-release pins allow for easy and fast exchange of the multiple locomotion systems. The robot is equipped with six sets of planetary reduction, MCD EPOS motors from Maxon Motors that have encoders and drives already integrated, 2 Bren-Tronics militarystandard batteries, Intel NUC onboard computer, Ubiquiti Rocket radio for long-distance wireless communication, and a pair of Axis cameras with high brightness LED lighting (Figure 3). Common configurations of the robot also allows for an instrumentation tower with Intel RealSense D435i and Authorized licensed use limited to: Carleton University. Downloaded on May 29,2021 at 22:41:33 UTC from IEEE Xplore. Restrictions apply. T265 cameras for odometry and mapping, an X-Sens Inertial Measurement Unit (IMU), an Ouster OS1-16 LIDAR, and a Tether Fathom-X from Blue Robotics for cable communication. This section describes the methodology and the preparation of the robot models to inspect a simulated confined environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000345_978-94-007-7194-9_2-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000345_978-94-007-7194-9_2-1-Figure2-1.png",
        "caption": "Fig. 2 Singular configurations: (a) elbow/knee singularities in both arms/legs (unavoidable); (b) wrist/shoulder singularity in 7-DoF right arm and hip singularity in right leg; (c) shoulder singularity in 7-DoF right arm (avoidable); (d) shoulder singularity in 6-DoF right arm (unavoidable)",
        "texts": [
            " Singular configurations and kinematically redundant limbs will be discussed in further detail in the following sections. The formulation of the forward differential kinematic problem in terms of accelerations can be obtained by differentiating (3) w.r.t. time: PV D J . / R C PJ . / P : (13) The solution to the respective inverse problem is then represented as: R D J . / 1 PV PJ . / P : (14) At a singular configuration, the end-link loses mobility, i.e., its ability of instantaneous motion in one or more directions. Figure 2a shows a robot standing upright with stretched legs and arms hanging at the sides of the body. This posture, though commonly used by humans, is usually avoided with a humanoid robot. The reason is that all four limbs are in singular configurations w.r.t. their root frames (fT g and fBg for arms and legs, respectively). Indeed, since the arms are fully extended, the hands cannot move in the downward direction w.r.t. the fT g frame. Similarly, since the legs are stretched, the fBg frame cannot be moved in the upward direction",
            " These singular configurations of the arms/legs are called elbow/knee singularities, respectively. They are characterized as unavoidable singular configurations: there are no alternative nonsingular configurations that would place the end-links at the same locations, i.e., at the workspace boundaries of each limb. As apparent from the example, unavoidable singular configurations are inherent to both redundant and nonredundant limbs (the arms and legs, respectively). Another singular configuration of the robot is shown in Fig. 2b. In addition to the elbow/knee singularities in the left arm/leg, there are singularities associated with the three-DoF spherical-joint type subchains in the shoulder and wrist joints of the right arm and the hip joint of the right leg, respectively. As apparent from the figure, two of the three joints axes in each of these joints are aligned which leads to mobility loss in one direction. The singular configuration of the arm is called wrist/shoulder singularity [8], while that for the leg, hip singularity. Figure 2c shows another type of singular configuration for the right arm. The singularity is due to the alignment of the two axes in the shoulder joint and the elbow joint being at 90\u0131. The end-link loses a translational mobility in the direction of the lower-arm link. This configuration is called shoulder singularity. Note that the end-link is placed within the workspace. In this case, the self-motion of the arm, i.e., a motion whereby the end-link is fixed (cf. Sect. 4.1), yields a transition to a nonsingular configuration, as shown in the figure. Such types of singular configurations are characterized as avoidable. A singular configuration of a kinematically nonredundant arm is shown in Fig. 2d. It occurs when the wrist center is placed on the axis of the first (root) joint at the shoulder. The translational mobility of the end-link is then constrained within the arm plane. This type of configuration is referred to as shoulder singularity. Confusion with the shoulder singularity of the redundant arm should be avoided. The singular configurations discussed so far are characterized by loss of end-link mobility in a single direction. As already noted, there are singular configurations when the end-link can lose mobility in more than one direction. This happens, for example, when the hip and knee singularities in the leg or the wrist/shoulder and elbow singularities in the arm occur simultaneously. Other combinations, as well as other types of singular configurations, do exist. Readers interested in in-depth analysis are referred to [8]. When a limb is in a singular configuration, the respective Jacobian matrix becomes rank deficient, reflecting the loss of mobility. In Fig. 2a, for example, all four limb Jacobians are rank deficient: rank J k. s/ D 5; k 2 fFr ; Fl ;Hr ;Hlg. The rank is five since each limb has lost mobility in one direction. Multiple singularities, e.g., an instantaneous elbow and wrist singularity, lead to further decrease in the rank of the Jacobian. Rank deficiency implies that the inverse kinematic solutions cannot be found with the formulas introduced in the previous subsections. The problem can be alleviated in a straightforward manner by avoiding singular configurations, e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003083_012017-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003083_012017-Figure1-1.png",
        "caption": "Fig. 1. Scheme of forces acting on the trailer-truck: X, Y \u2013 coordinate axes associated with the truck center of mass; \u0421\u0422, \u0421P \u2013 centers of mass of the truck and semi-trailer; S \u2013 the fifth wheel center; scl \u2013 distance from the center of mass of the truck to the fifth wheel center; \u00a0 \u0430j \u2013 acceleration of the center of mass of the truck; \u0430V \u2013 linear velocity of center of mass of the truck; z \u2013 the angular velocity of rotation of the truck around a vertical axis passing through its center of mass; scP \u2013 force acting at the fifth wheel center from the side of the semi-trailer to the truck during braking; 1 1,\u00a0l rX X \u2014 longitudinal forces acting on the left and right wheels of the first axle of the truck from the road; 1l 1r\u0398 ,\u00a0\u0398 \u00a0 \u2013 steering angles of left and right wheels of the first axle of the truck; 1l 1r,\u00a0  \u2013 slip angles of left and right wheels of the first axle of the truck; 2l 2r,\u00a0  \u2013 slip angles of left and right wheels of the second axle of the truck; \u03b3 \u2013 jackknifing angle of the trailer-truck",
        "texts": [
            " These facts allows to apply a corrective change in the angle of rotation of the steered wheels (steering), which will help to keep the truck on the path set by the driver. A simplified mathematical model of trailer-truck planar motion will be consider to compose the equations of trailer-truck motion. The model has several assumptions: a) the slip angles of right and left wheels of each axis are the same; b) the steering angles of the wheels and wheel slip angles are small, i.e. do not exceed 10\u00b0; c) slip resistance coefficients for all wheels of the axle are the same. The design diagram of the forces acting on the trailer-truck in curved motion is shown in Fig. 1. Design Technologies for Wheeled and Tracked Vehicles (MMBC) 2019 IOP Conf. Series: Materials Science and Engineering 820 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/820/1/012017 Consider the movement of a truck in braking mode, replacing the effect of the semi-trailer with the force scP applied to the fifth wheel center on the truck, acting along the longitudinal axis of the semitrailer For a rear drive two-axle vehicle the following differential equations were obtained in [21]: ( ) 2 2 1 1 2 1 1 2\u0398 4 4 y ya sr a a z a a z K KV g L g L L V G J V G J      = +  \u2212  \u2212 +  \u2212 \u2212  +        ( ) 2 1 1 1 1 \u0398 \u0398 ; 2 4 a sr sr y z a a a z a a a z j g L g L P M X V G V J V V G J   + \u2212  + + \u2212 +    ( ) 2 2 2 1 2 1 1 2\u0398 4 4 y ya sr a a z a a z K KV g L g L L V G J V G J      = +  \u2212  \u2212 \u2212  \u2212 +  \u2212        2 1 2 1 \u0398 ; 2 4 a sr y z a a a z a a a z j g L g L P M X V G V J V V G J   \u2212  + \u2212 \u2212 \u2212    ( ) ( )sc 1 2 1 sin ;\u00a0 ;\u00a0 sin \u0398 z z sc a y sc \u0441\u0440 L M P l V P P  =  = =  +  \u2212  ( ) ( ) 2 2 sc sc 1 1 1 1 sin sin \u0398 , 2 2 y a a A y A y sr a z a z z z z a z a K L j L j P l P l X V J V J J V J V      = \u2212 + +   \u2212  \u2212 +         (1) where 2 1A =  \u2212  \u2013 difference between rear and front slip angels; 1 1l 2rX X X= + \u2013 full force on the front axle of the truck; yK \u2013 total (for the vehicle axis) coefficient of resistance to tire retraction; 1 1 1 \u0398 \u00a0\u0398 \u0398 2 l r sr + = \u2013 average steering angle; l \u00a0 ,\u00a0 1,\u00a02 2 i ir i i  +  = = \u2013 average slip angle of i-truck axle; aG \u2013 truck weight; L \u2013 truck wheelbase; zJ \u2013 the moment of inertia of the truck relative to the vertical axis passing through its center of mass",
            " The direct Lyapunov method reduces the problem of studying the stability of system (3) to studying the properties of the function V(X) and its first derivative, which is calculated by the formula ( ) ( ) 2 2 sc 1 2 sin 2 y a yA a z a z z KdV X V L j AX P l dt X V J V J   = = \u2212  + +      (5) To find the optimal control vector U(t), we use the method of optimal damping of transients V.I. Zubova [22]: ( ) ( )max V U t U sign R t X   = \u2212     (6) The magnitude order of the summands of equation (5) and their signs based on the positive directions of vectors adopted in Figure 1 are shown in table 1. and the function ( )V X is a Lyapunov function. Then the optimal control (6) takes the form ( )   ( ) 2 sc 1 2 1 sin 2 max A a y z z a z a U t U sign D L j D P l X J V J V = \u2212 \u2212    =  \u2212 +    (7) Because 2 1 0,\u00a0then\u00a0 0, 2 a z L X D V J   and expression (7) can be written ( ) ( )max AU t U sign= \u2212  (8) Design Technologies for Wheeled and Tracked Vehicles (MMBC) 2019 IOP Conf. Series: Materials Science and Engineering 820 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/820/1/012017 A control restriction is introduced ( ) ( ) 0 ,maxU t t=  =  \u2212  where ( )t \u2013 actual hitch angle, 0 \u2013 value of hitch angle when the driver presses the brake pedal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003635_s00170-020-06321-7-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003635_s00170-020-06321-7-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of the contact between the abrasive particle and the honing surface",
        "texts": [
            " Under the current conditions, the cylindricity of the honed cylinder liner can be improved by selecting the optimized honing processing parameters. During the honing process, the convex abrasive particles on the surface of the honing oilstone contact with the cylinder liner surface. Oilstone abrasive particles are randomly distributed in the honing oilstone. Because the cylindricity is macroscopic, the abrasive particles of the oilstone can be regarded as spherical, and the distribution of the abrasive particles is assumed as uniform. The abrasive particles contact with the processed surface to produce the cutting depth, as shown in Fig. 2: The average cutting depth h can be calculated by the Brinell hardness model [23]: HB \u00bc P As \u00f01\u00de where P is the average honing pressure, and As is the contact area of a single abrasive particle with the honing surface. As \u00bc 2\u03c0rh \u00f02\u00de where r is the radius of the abrasive particle, and h is the average cutting depth of a single abrasive particle. Hence, the average cutting depth of a single abrasive particle is h \u00bc P 2\u03c0rHB \u00f03\u00de The honing pressure [24] can be calculated by P \u00bc P 0 Ap Atan\u03b1 \u00f04\u00de where P is honing pressure (105 Pa), P\u2032 is oil pressure of the honing machine (105 Pa), Ap is piston area of honing head (cm2), A is the total area of oilstone (cm2), and \u03b1 is generatrix slope of the feed cone (\u00b0)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002248_icca.2016.7505337-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002248_icca.2016.7505337-Figure2-1.png",
        "caption": "Fig. 2 Topologies",
        "texts": [
            " t h t t t te Re Defining ( ) [ ( ), ( ( ))],t t t ty e e then ( ) ( ) ( )V t t ty y , where is defined in (15). If the equation (15) is satisfied, then the system (14) is asymptotically stable, i.e., the heterogeneous multi-agent system (12) reaches the consensus asymptotically. V. SIMULATION AND RESULTS In this section, the gravity acceleration is assumed to be 10g for the quadrotor model. A. Consensus Problem without Time Delay of System 1) The heterogeneous multi-agent system (9) consists of six agents (see the topology 1G in Fig. 2 (a)). In 1G , the solid and hollow circles represent the fourth-order (5) and first-order (6) integrators respectively. The coupling weights in 1G is chosen as 1, and the control parameters in (7) are: 1 2 3 41( 1), 7, 9, 5k k k k k , so Prerequisite 1 holds. Left eigenvector corresponding to the zero eigenvalue of H in (9) is = 3 12[2,4,2,4,2,4,1,2,0,1,1,1]l , which satisfies 1 2 3 =1l 1 . Choosing 1(0) 16,x 2 (0) 3,x 3 3 3 3(0) (0) (0) (0) 5x v p w , 4 (0) 7,x 5 5(0) (0)x v 5 5(0) (0) 1p w , 6 (0) 10,x (0) [5,1,10,2,5 15,g 3,20 20,4 4,16,3,7,10] .g g g From Theorem 1, the system (9) achieves an asymptotic consensus, lim ( )t ix t 23, 1, 6i and lim ( ) lim ( )t i t iv t p t lim ( ) 0, 3,5t iw t i , which are the same as the simulation results shown in Fig.3 and Fig.4. 2) Assume that the topology of heterogeneous multi-agent systems (5) and (6) is a balanced and connected graph 2G shown in Fig.2 (b). The agents initial states are the same as that in 1), and the control parameters satisfying Remark 1 are chosen as 1 2 31( 1), 4, 6,k k k k 4 4k for algorithm (7). From Remark 1, the system (9) reaches a quasi-average consensus asymptotically, and *lim ( ) 33t ix t , 1, ,6i , lim ( ) lim ( ) lim ( ) 0t i t i t iv t p t w t , 3,5i (see Fig.5 and Fig.6). B. Consensus Problem with Time-varying Delay The topology 1G of heterogeneous multi-agent systems (10) and (11) is shown in Fig. 2 (a), and the coupling weights in 1G is 1. The agents initial states are randomly generated. Choose 1 2 31( 1), 2, 4k k k k , 4 3k in the algorithm (10), and we get 0.333h from Theorem 2. The input time-varying delay is chosen as ( ) 0.333 sin( (1 ))t t , and the system (12) reaches a consensus asymptotically. The simulations are shown in Fig.7 and Fig.8. VI. CONCLUSIONS By linearizing the differential driven unmanned ground vehicles (UGVs) and quadrotor unmanned aerial vehicles (UAVs) as first-order integrators and fourth-order integrators respectively, we investigates the consensus problem for a class of heterogeneous multi-agent systems composed of first-order integrators and fourth-order integrators in this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002487_s2010132517500018-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002487_s2010132517500018-Figure1-1.png",
        "caption": "Fig. 1. (a) Detector of e\u00aeort, (b) dual detector of e\u00aeort.",
        "texts": [],
        "surrounding_texts": [
            "The pseudo BG model consists of describing the thermal transfers in a household refrigerator compartment, which is exposed to an outdoor natural cold air\u00b0ow. This latter is spread out in a cavity covering the side wall of the appliance and modeled by a modulated \u00b0ow source (MSf) using the outdoor temperature ( 20 C). It characterizes the frosty climate in Nordic countries. From the outside environment, the cold air\u00b0ow is delivered via a small fan and circulated in an inlet duct, which is connected to the cavity through an inlet opening. Another one linked the cavity to the outlet duct, from which the 1750001-3 In t. J. A ir -C on d. R ef . 2 01 7. 25 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by U N IV E R SI T Y O F N E W E N G L A N D o n 04 /0 7/ 17 . F or p er so na l u se o nl y. cold air\u00b0ow will be evacuated towards the outside, as shown in Fig. 4. 3.1. Word bond graph model It is presented in Fig. 5. The fan is used to push the cold air\u00b0ow stemming from the outside environment through the inlet duct. It is modeled by a modulated \u00b0ow source (MSf). The \\Refrigerator\" block presents the word BGmodel of the thermal transfers in the refrigerator compartment without the use of any second cooling source (Aridhi et al.15). The thermal transfers between the refrigerator and the fan are made by forced convection. 3.2. Causal pseudo bond graph model The pseudo BG model highlights the system components, the sources, the thermal variables (temperature and heat \u00b0ow) and the direction of the energy distribution. It is developed using 20-sim software, presented and depicted in Figs. 6 and 7, respectively. The ambient temperature Ta is modeled by an e\u00aeort source (Se:Ta). The cold air\u00b0ow is modeled by a \u00b0ow source (MSf) that is modulated by the outdoor temperature Tout (a square signal varying between 20 C and 10 C). It is modeled by a modulated e\u00aeort source (MSe:Tout). The inlet duct has the pyramidal form and is made of plastic material. The width of the big and small bases is 0.2m and 0.1m, respectively. The thickness of its walls is 3mm. It is connected to the cavity through an opening, allowing then the distribution of the outdoor cold air\u00b0ow in the installation. It is modeled by the \\Duct in\" block, which describes the thermal losses at the inlet opening level. They are modeled by the R element (R5). The thermal losses at the level of the inlet duct walls Fig. 4. Outdoor cold air\u00b0ow distribution at the side wall level of the refrigerator compartment. 1750001-4 In t. J. A ir -C on d. R ef . 2 01 7. 25 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by U N IV E R SI T Y O F N E W E N G L A N D o n 04 /0 7/ 17 . F or p er so na l u se o nl y. with the ambient air and the outside air are respectively modeled by the R elements (R11) and (R6). However, the heat storage in the inlet duct is modeled by the C element (C5). The outlet duct is rectangular in shape, 3m length and has the same thickness of the inlet duct walls. It allows the evacuation of the air towards the outside environment. The \\Duct out\" block is the BG model of thermal transfers involved in the outlet duct. It describes the thermal losses at the outlet opening level, which are modeled by the R element (R8). Furthermore, it highlights the thermal losses at the level of the outlet duct walls with the ambient air and the outside environment. They are respectively modeled by the R elements (R9) and (R10). In addition, the heat storage is modeled by the C element (C6). The evaporator temperature source (Mse: Tevap) is modulated by the temperature at the evaporator wall level according to the compressor speed."
        ]
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure7.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure7.2-1.png",
        "caption": "Fig. 7.2 Pictorial diagram of solid particle erosion test rig",
        "texts": [
            " The tooling cost and the time required to fill the mould is very less as compared to the other fabrication process. The complete fabrication process is performed at room temperature. Concentration of the fibers and the adhesion between the layers is improved by the application of the pressure. The flowchart of the different processes involved in the fabrication process of composites using VARTM process is shown in Fig. 7.1. The solid particle erosion tests were performed using Ducom TR 471 according to ASTM D76 (Sharma et al. 2018). The pictorial diagram of solid particle erosion test rig is shown in Fig. 7.2. The solid particles are fed from hopper and themass flow rate is controlled by the conveyor belt. Themix is carried by the conveyor belt and through the nozzle of diameter 1.5 mm made to strike the composite sample with desired velocity. The velocity of the erodent is selected in the range of 30\u201360 m/s. The standoff distance between the sample and the nozzle is kept 10 mm. The impingement angle is selected as 45, 60, 75 and 90\u00b0. Silica is used as a erodent and the discharge is maintained at constant rate i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002061_1.4033362-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002061_1.4033362-Figure4-1.png",
        "caption": "Fig. 4 Solid-element finite element model of crankshaft",
        "texts": [
            "org/pdfaccess.ashx?url=/data/journals/jotre9/935662/ on 03/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use @ @h h3 @p @h \u00fe R2 @ @z h3 @p @z \u00bc 12gR2 _e cos h /\u00f0 \u00de\u00bd \u00fee _/ x 2 sin h /\u00f0 \u00de (6) where g is the dynamic viscosity of lubricant, x is the angular speed of crankshaft. 2.2.4 Model of Crankshaft Dynamic Stress. The dynamic stress of crankshaft is calculated by the finite element method. The solid-element finite element model of crankshaft is established in ANSYS software (shown in Fig. 4), which consists of 39,950 elements and 72,462 nodes. The mesh generation on the journal surface of crankshaft is controlled to be in conformity with that used in the lubrication analysis of bearing by the finite difference method. The dynamic stress of crankshaft is calculated by applying the dynamic nodal force on the journal surface of crankshaft. The dynamic nodal force is gained by the transformation of the oil film pressure of bearing based on the following equation: Fxi t\u00f0 \u00de \u00bc \u00f0zi\u00fe1 zi \u00f0hi\u00fe1 hi 1 4 p hi; zi; t\u00f0 \u00de \u00fe p hi\u00fe1; zi; t\u00f0 \u00de \u00fe p hi; zi\u00fe1; t\u00f0 \u00de \u00fe p hi\u00fe1; zi\u00fe1; t\u00f0 \u00de R cos hdhdz Fyi t\u00f0 \u00de \u00bc \u00f0zi\u00fe1 zi \u00f0hi\u00fe1 hi 1 4 p hi; zi; t\u00f0 \u00de \u00fe p hi\u00fe1; zi; t\u00f0 \u00de \u00fe p hi; zi\u00fe1; t\u00f0 \u00de \u00fe p hi\u00fe1; zi\u00fe1; t\u00f0 \u00de R sin hdhdz 8>>< >>: (7) where Fxi and Fyi are the dynamic nodal force components of i node, hi is the angular coordinate of i node, zi is the axial coordinate of i node, p(hi, zi, t) is the oil film pressure of bearing at t moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002417_s10404-016-1811-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002417_s10404-016-1811-5-Figure1-1.png",
        "caption": "Fig. 1 Design and assembly of membrane-free \u03bcMFC. a Three types of \u03bcMFC chip designs for comparison: (1) Y-shape; (2) expanded cathode and anode chambers without micro-stoppers; and (3) microstoppers in the expanded anode chamber. b Schematic illustration of the major components of \u03bcMFC. c The image of an assembled \u03bcMFC. The gold electrodes were connected to a digital multimeter for OCV measurement. The enlarged inset shows the laminar co-flow in the downstream microchannel after electrolytes are pumped in. The yellow solution on the left is catholyte (color figure online)",
        "texts": [
            " Sodium lactate, NH4Cl, KCl, Na2HPO4, Na2SO4, MgSO4\u00b77H2O, piperazine-N,N\u2032-bis(2-ethanesulfonic acid) (PIPES), NaCl, CaCl2, Na2MoO4, CuCl2\u00b72H2O, FeCl2\u00b74H2O, MnCl2\u00b74H2O, CoCl2\u00b74H2O, ZnCl2, H3BO3, NiSO4\u00b76H2O, Na2SeO3\u00b75H2O, Na2WO4\u00b72H2O and riboflavin were obtained from Sigma-Aldrich. Deionized (DI) water was collected from Millipore Synthesis A10 (Molsheim, France). Shewanella oneidensis MR-1 (ATCC 700550) was acquired from American Type Culture Collection (Manassas, VA), and Escherichia coli DH5\u03b1 (DSM 6897) was obtained from the German Collection of Microorganisms and Cell Cultures (DSMZ) (Braunschweig, Germany). Three types of \u03bcMFC chips were designed to compare their performance and identify the optimal design (Fig. 1a). The first one has a simple Y-shape. The second is with expanded cathode and anode chambers. The third is with micro-stoppers in the anode chamber. The purpose of including the expanded chambers and stoppers is to maximize the bacteria retention on the anode and increase the surface area of electrode for efficient electron transfer. A glass substrate (1.5 in. \u00d7 3 in.) was used to pattern the electrodes. A layer of negative photoresist (SU8-3010) was spin-coated on the glass substrate and then exposed to UV irradiation through a film photomask",
            " A mixture of silicon elastomer base and curing agent at a mass ratio of 10:1 was cast onto the SU-8 mold and cured at 70 \u00b0C for 3 h. Then, the PDMS slab with designed channel structure was sliced and peeled off from the mold. Before assembling the parts, the PDMS chip and gold-patterned glass substrate were rinsed with DI water and autoclaved at 120 \u00b0C under 10 atm. for 20 min for sterilization. The PDMS chip was manually stacked onto the substrate with the microchannel aligned along the gap between two electrodes (Fig. 1c). A dual-syringe pump (KDS Legato 180) was used to infuse the anolyte and catholyte into the micro-chambers. A digital multimeter (3146A ESCORT) was used to measure the open-circuit voltage (OCV) produced by the \u03bcMFC. The gold electrodes were connected to the digital multimeter by electrical clips. An aluminum foil was placed between the gold electrode and the electrical clip to ensure close contact for electric conduction. Data were recorded at 2-min interval for totally 1\u20131.5 h until a stabilized output was achieved",
            " Although the micro-stoppers did not change the surface area of the electrode significantly, they increased the flow resistance and decreased the flow velocity in the anode chamber as shown by numerical simulation (Fig. S1, Supporting Information). As reported in the literature on laminar flow-based micro fuel cells (Choban et al. 2004; Sprague et al. 2009), the change of anolyte flow patterns Microfluid Nanofluid (2016) 20:144 1 3 Page 5 of 8 144 could affect the \u00b5MFC internal resistance to proton transport and hence the OCVs. The laminar co-flow was developed immediately after the catholyte and bacteria-laden anolyte were introduced into the \u03bcMFC (Fig. 1c, inset). The measured OCV quickly adjusted and increased to a stabilized output (Fig. 2a). The time required to reach a stabilized output ranged from about 10 to 90 min, depending on the specific chip design and bacterial species. The \u03bcMFC with expanded reaction chambers typically required a longer time for output stabilization. In contrast, it usually takes a few days for conventional MFCs to reach stabilized output (Fan et al. 2012; Yu et al. 2011), while a few hours for many existing miniaturized MFCs (Qian et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002607_yac.2016.7804934-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002607_yac.2016.7804934-Figure5-1.png",
        "caption": "Fig. 5. Attitude angle definition in fixed-wing mode.",
        "texts": [],
        "surrounding_texts": [
            "The aircraft structure diagram is shown in Fig. 3, and the real product is shown in Fig. 4. The fixed-wing fuselage\u2019s layout is blended wing body layout, on the top of the fuselage two brushless motors are installed for providing the hover lift in rotary-wing mode and pull in fixed-wing mode, at the bottom of the fuselage two control surfaces are installed for controlling the vehicle attitude. To use the rotary-wing mode as a reference, the establishment of the body coordinate system F: OXYZ, the original point is the aircraft\u2019s center of gravity, OX axis perpendicular to the wing plane pointing to forward (perpendicular to the paper inwards), OY axis pointing to right in wing plane, OZ axis in wing plane perpendicular to the plane OXY pointing down. The vehicle mounted two brushless motors side by side in the top of the fuselage, drives a pair of positive and negative propellers to provide an upward pull, balance each other their respective reaction torque, via motors differential rotation to achieve the body rotate around the OX axis (roll, right and left movement), define the right roll as the positive direction of roll angle; in addition, the surfaces at bottom of the body are two independent controlled control channels, two differential rudder deflection about OX axis\u2019s positive and negative directions can achieve body rotate around OZ axis (yaw control), define the right yaw as the positive direction of yaw angle; two simultaneous rudder deflection about OX axis\u2019s positive or negative direction can achieve body rotate around OY axis (pitch, front and back movement), define backward flight as the positive direction of a pitch angle. A. Attitude Angle Definition in Rotary-wing Mode (Fig. 3) Pitch angle 1: the angle between the OX axis and the horizontal plane, backward flight as the positive, range (-90 \u00b0, 90 \u00b0); Yaw angle 1: the angle between north and the projection in horizontal of OX axis, projection in horizontal plane at the north east is positive, range [-180 \u00b0, 180 \u00b0]; Roll angle 1: the angle between the horizontal plane and OY axis, the right motor sink, the left motor raise is positive, the range of (-90 \u00b0, 90 \u00b0). The body rotate 90 degrees around the OY axis in the negative direction, the aircraft will transit into fixed-wing mode, the principles of body rotate around the three axis in fixed-wing mode is same as the rotary-wing mode, however, relative to traditional fixed-wing aircraft, roll angle of rotary-wing becomes yaw angle of fixed-wing (change nose pointing in airplane mode), yaw angle of rotary-wing becomes the roll angle of fixed-wing (and opposite), pitch angle in the two mode are similar, they are both rotate around OY axis (but there is a difference of 90 degrees), the motors provide lift in rotary-wing mode while provide pull for forward movement in the fixed-wing mode, the lift of the fixed-wing is produced by the relative motion between are and the wing."
        ]
    },
    {
        "image_filename": "designv11_30_0001143_acc.2013.6579933-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001143_acc.2013.6579933-Figure3-1.png",
        "caption": "Figure 3. Blade Lead-Lagging Motion",
        "texts": [],
        "surrounding_texts": [
            "Modeling is based on two key ideas. First, physics principles and appropriate assumptions are used to directly result in nonlinear ordinary differential equations (ODEs), without resorting to partial differential equations models. Second, flight and blade dynamics modes that are critical for safe and performant helicopter operation are captured. Multibody dynamics was used to include all helicopter components: fuselage, articulated main rotor with 4 blades, empennage, landing gear, tail rotor [4]. The key modeling steps are described next. 978-1-4799-0178-4/$31.00 \u00a92013 AACC 794"
        ]
    },
    {
        "image_filename": "designv11_30_0002387_cae.21769-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002387_cae.21769-Figure1-1.png",
        "caption": "Figure 1 Quadcopter diagram with coordinate frames. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary. com].",
        "texts": [
            " Fuzzy logic is also a nonlinear controller, a feature that makes it suitable for highly nonlinear systems like quadcopters. The nonlinear dynamic equations of motion for the six degrees of freedom of a quadcopter have been derived from the first principle. Noise is simulated in this platform and is applied to the position and angle measurements. To reduce the noise impact on the behavior of the quadcopter, a Kalman filter has been used. Two coordinate frames are used: a body-fixed coordinate frame or local frame {xb, yb, zb} and a global reference frame {xn, yn, zn} as depicted in Figure 1. This global reference frame is assumed to be inertial for two reasons: (1) The earth completes one rotation per day which is a low speed relative to its mass and size. (2) The quadcopter does not fly for long periods of time and large distances. The following assumptions are considered while developing the mathematical model: (1) The body frame coincides with the center of mass of the quadcopter. (2) The quadcopter structure is symmetrical about each one of its local coordinates. (3) The quadcopter structure is rigid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001589_saci.2013.6608956-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001589_saci.2013.6608956-Figure1-1.png",
        "caption": "Figure 1. DC current decay test at standstill for dq inductance identifications",
        "texts": [
            "5p\u03bbPMiq, and the reluctance torque driven by the difference between Ld and Lq. For SPMSM, the magnetic torque is available only because Ld=Lq=Ls, and therefore the reluctance torque is equal to zero. \u2013 147 \u2013 978-1-4673-6400-3/13/$31.00 \u00a92013 IEEE Resistance Calculation The dq axis inductances of the SPMSM tested prototype are determined after performing dc current decay tests at standstill, with the rotor aligned along d axis, and along q axis, respectively. The electrical diagram of the test is illustrated in Fig. 1. The current decay occurs in the phases B and C, which are actually connected in series, while in the phase A different value of constant current (icc) is injected. The dq inductances are calculated with the expression: , (2 ) / (2 )= \u22c5 \u22c5 + \u22c5\u222b \u222bd q s D contL R idt V dt I , (7) where: Rs - phase resistance; i - decay current, VD - diode voltage drop, Icont - the direct current before the decay. The dependencies (\u03bbd - \u03bbPM) =Ldid and \u03bbq =Lqiq are illustrated in Fig. 2 a,b. The cross coupling is visible for both axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure5-1.png",
        "caption": "Fig. 5 Analysis of in series connected two-carrier planetary gear train",
        "texts": [
            " The most commonly used version is the first with two compound shafts and three external shafts (Fig. 4a). As seen from the figure, one of the compound shafts has no outset\u2014it is an internal compound shaft, so that at both ends two equal sized but opposite direction torques are acting. The external torque of the other one, the external compound shaft, which has outset, equals the sum of external torques of the two coupling shafts. The mentioned things above are illustrated by three examples: \u2022 in series connected gear trains (Fig. 5); \u2022 a closed differential gear train with internal division of power (Figs. 6 and 7); \u2022 a closed differential gear train with internal circulation of power (Figs. 8 and 9); Moreover, these are the three scenarios in this way of linking the two planetary gear trains\u2014with two compound shafts and three external shafts. Other cases are not possible. The first case (Fig. 5) is undoubtedly the simplest and most frequently used. There the whole power passes consecutively through two elementary component gear trainss. In both component trains the relative powers Prel I and Prel II are transmitted from the sun gears 1 and 4 of the ring gears 3 and 6. 0,75 3 0, 25 circ A P P \u2212= = 0, 25 4 3,75 0i A B CT T T T= + + = \u2212 + =\u2211 4 16 0,25 B A T i T \u2212= \u2212 = \u2212 = + + 4 3I IIt t= > = SI SII 0BP <0AP > I II rel IP rel IIP 1 2 3 5 6 4 rel IIP BP AP rel IP4It = 3 1 0,75 I T t = = \u2212 4 1= + 3 6 3T= \u2212 = \u2212 1 4 0, 25AT = + = + ( )4 6 4 B ST T T \u2261 = = \u2212 + = = \u2212 1 2 3 4 5 6 7 3IIt = ( )1 3 3,75 T T \u2261 = = \u2212 + = = + 6 4 3 IIt T= \u22c5 = = + circP min max; ;A D C D BT T T T T T\u03a3\u2261 \u2261 \u2261 3 1 1,33",
            " For example, if in the first component train is begining from the real torque T 01 of the sun gear 1, for the size of the real torque T 03 of the ring gear 3 there are two options: T 03 \u00bc tI T 01 g0I in the direction of the relative power Prel I from the sun gear 1 to the ring gear 3; T 03 \u00bc tI T 01 g0I in the direction of Prel I from the ring gear 3 to the sun gear 1; In these formulas g0 I is internal efficiency of the I component gear train, i.e. in transmission of power from the sun gear 1 to the ring gear 3, with fixed carrier SI xSI \u00bc 0\u00f0 \u00de: What was said up to now is true for the second component gear Fig. 5 shows: \u2022 determination of ideal and real torques; \u2022 check on the equilibrium of the torques, ideal and real, i.e. their correct calculation; \u2022 determination of speed ratio ik; \u2022 determination of torque ratio iT (torque transformation); \u2022 determination of efficiency g: The next two cases of closed differential gear trains are more interesting because in one case (Figs. 6 and 7) there is an internal division of power, which is beneficial and in the other case (Figs. 8 and 9)\u2014internal circulation of power, which is not very beneficial"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000175_icems.2019.8922428-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000175_icems.2019.8922428-Figure6-1.png",
        "caption": "Fig. 6. Positioning of the first circle",
        "texts": [
            " Since the first circle cannot be set to the coordinates x = y = 0 due to the angle \u03b1 of the parallelogram, the length x must be considered for the displacement of the circle. First the height y at which the circle contacts the bottom of the slot is calculated. The length x1 can then be calculated using the angle functions and the Pythagorean theorem. According to the above definition, however, the length x is required for positioning the circle. The calculation of the length x, by which the wire is shifted in x-direction, is only valid for the 1st layer. The angular relations for the calculation of the length x are illustrated in Fig. 6. After positioning a single circle, the system checks whether the corresponding circle is completely within the area of the parallelogram in both spatial directions. The positions of the circle or circular path are queried using several points and compared with the identical points of the parallelogram. The further circles in x-direction are shifted by the respective diameter. If a circle does not fulfill the described conditions, it must be removed again. Thus the 1st layer of the parallelogram for the corresponding circle diameter is completely filled"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000033_aim.2019.8868339-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000033_aim.2019.8868339-Figure3-1.png",
        "caption": "Figure 3. The importance of precision during obstacle detection.",
        "texts": [
            ", the circular arc corresponding to a given angle increases with the range and any error can easily significate a difference of several centimeters, a major drawback in ultra-long range applications1. The two following examples will allow to better understand the importance of precision in leveling. 1 At 100 m, 0.1\u00b0 corresponds to 17 cm height. First, let\u2019s suppose someone is driving at 100 km/h. The reaction distance, that is, the distance traveled by the vehicle during the time an obstacle is detected and the brake is activated, is around 40m. The stopping distance for the same value of speed is around 100m. In Figure 3, \ud835\udf03 is the angle formed by the vertical \u210e and the cut-off line that lies at a distance \ud835\udc65 from the front of the vehicle. We can easily deduce that, for a given value of angular precision \ud835\udeff and a stopping distance \ud835\udc652 \ud835\udc65 = \u210e ( \ud835\udc652 \u2212 \u210e tan \ud835\udeff \u210e + \ud835\udc652 tan \ud835\udeff ) Classic levelers have a precision in the range of \u00b10.1\u00b0. Under the former conditions, \ud835\udc65 corresponds to 80 m. So additional 20 m are needed to perceive the obstacle and the minimum 100 m needed to stop the vehicle, are no longer suitable, ultimately requiring more effort on the brakes to avoid collision"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000380_iros.2011.6095187-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000380_iros.2011.6095187-Figure6-1.png",
        "caption": "Fig. 6. Mechanism of the window cleaning robot",
        "texts": [
            " (7) If the distances between the robot and window flames Lx, Ly are impossible to measure because window frames are not fully-flat or the heights of window frames are low, the position information of ceiling, floor or wall can be used. A window cleaning robot can follow the horizontal parallel motion trajectory exactly with the trajectory tracking control method proposed in the section II. However, the robot can not assure a complete cleaning of a glass window. Therefore, this section proposes a novel dirt detect sensor for a window cleaning robot shown in Fig.5 and a motion control method to guarantee completeness of window cleaning by using the dirt detect sensor. Mechanism of the window cleaning robot is shown in Fig.6. Traditionally, dirt on glass window is detected and evaluated by human subjectively. A quantitative dirt detection method by measuring amount of transmitted light which decreases in dirty place has been proposed [9]. In this method, a light emitting section and a light receiving one are necessary to face each other at a constant distance, because amount of light is in inverse proportion to a distance from an illuminant. However, the receiving section cannot face the emitting one constantly on the developed robot for window cleaning because stick-slip vibration is generated by friction resistance of the cleaning unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000039_aim.2019.8868900-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000039_aim.2019.8868900-Figure10-1.png",
        "caption": "Figure 10. Experimental setup that consists of linear guide driven by ball screw with AC motor, grip with force sensor, and additional mass.",
        "texts": [
            " Time responses of estimated external force \u02c6 df in comparison with DOB estimates in Case #2: Sinusoidal external force is adding in the absolute coordinate system while the manipulator is keeping a non-singular configuration. In this section, experimental conducted using a linear stage to evaluate the proposed external force estimator is described. To show the feasibility of the proposed method, one degree-of-freedom experimental setup was used this time, and the higher degree-of-freedom case will be demonstrated in the near future. The experimental setup consists of a linear guide with one degree of freedom driven by a ball screw driven (Fig. 10). The actuator is a torque-controlled AC motor. A DSP system operating at a sampling period of 5 ms is used to implement the control law. In the experiment, a mass of about 0.4 kg is added to the grip as shown in Fig. 10. Then, the force required to move the mass including the grip, Nf , is estimated by the RNN inverse dynamics model, and by subtracting the measured force by the sensor, f , from Nf , the estimate of the external force applied by a person, \u02c6 df , is estimated. Thus, in this translational motion, (27) and (28) are simplified as \u02c6 d Nf f f  . (29) As the experimental procedure, the linear guide drives the grip and the additional mass by tracking two kinds of reference trajectories of the position controller. One is a sin wave with an amplitude of 0.02 m and a period of 1 Hz, and the other is a band-limited M sequence with an amplitude of 0.02 by being passed through a low-pass filter with a cutoff frequency of 5 Hz. During the motion, the force required to move the mass including the grip is measured by the force sensor shown in Fig. 10. Fig. 11 shows the time response of the state when the linear guide is moved with these two kinds of trajectories, including the differential value estimated by the DOB type differentiator. This T N y y y   x     is the input (feature) vector of RNN. In the experiment, first, 4 sets of data without applying external force to the grip are measured and averaged. Using the averaged data of 4 sets, the RNN is trained so as to estimate the force required to move the mass including the grip as Nf from the motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003805_ipemc-ecceasia48364.2020.9368211-Figure1-1.png",
        "caption": "Fig. 1 Main transmission of servo press.",
        "texts": [
            " In order to solve the nonlinearities problem, the dynamical model of servo press is derived using computed torque module and PD parameters adaptively based on the change of angle error, speed error and reference acceleration in the second part. Thirdly, the PTC controller is designed for this model which is disturbed by unknown mathematical characteristics. Finally, through simulations, the proposed scheme shows valid control performance. II. DYNAMIC MODEL OF SERVO PRESS A press mechanism driven by an IM is shown in Fig.1 and Fig.2, where m1, m2, m3, mc, ms2 and ms are the masses of the three connecting rods, the crank, shaft2 and the slide, respectively. L1, L2 and L3 are the lengths of connecting rods, respectively. R is the radius of rotating crank. d and h are the horizontal and vertical distances from O to B. The slide can move along the guides up and down when the IM shaft rotates. The main task of this research is to track the desired motion of This study is supported by the \u201cKey Innovation Special Program\u201d of Qilu University of Technology"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000665_1.a33330-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000665_1.a33330-Figure6-1.png",
        "caption": "Fig. 6 Projectile side view and parameters.",
        "texts": [
            " The projectile dynamics are subject to the flight mechanics principles [12], which lead to a nonlinear parameter-dependent model. In the present case, this model is simplified thanks to the absence of linear motion and the following assumed hypotheses for the study of the pitch axis: the airspeed V is held constant and the other axes are locked at zero; therefore, p 0, r 0, \u03d5 0, and \u03c8 0, where \u03c9 p q r \u22a4 is the body rate vector and \u03d5 and \u03c8 denote, respectively, the body roll and yaw angles. Furthermore, in this experimental setup the altitude h is constant (ground level). Figure 6 presents the projectile in a pitching motion and the remaining parameters, which are the angle of attack (AoA) \u03b1, which is equal to the pitch angle \u03b8, the pitch rate q, and the control fin deflection \u03b4m. The pitch axis nonlinear model is linearized around a family of equilibrium (or trim) points, resulting in a second-order quasi-LPV model. To account for aerodynamic disturbances and measurement noise, a second-order disturbance model is appended to the projectile quasi-LPVmodel with the same poles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001575_1464419314540901-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001575_1464419314540901-Figure6-1.png",
        "caption": "Figure 6. Detailed cascade plots in the speed region of 0\u201330,000 r/min for Figure 5 respectively in the horizontal and vertical directions (e\u00bc 0.1 lm). f1 means the rotational frequency and f2 the VC frequency. Lxvc (nxvc\u00bc 4400 r/min, fxvc\u00bc 372.5 Hz), Lxvc2 (nxvc2\u00bc 14,200 r/min, fxvc2\u00bc 1202 Hz), Lyvc (nyvc\u00bc 6300 r/min, fyvc\u00bc 533.3 Hz), Lyvc2 (nyvc2\u00bc 20,800 r/min, fyvc2\u00bc 1761 Hz), Lxb (nxb\u00bc 10,500 r/min, fxb\u00bc 353.4 Hz), and Lyb (nyb\u00bc 10,500 r/min, fyb\u00bc 535.4 Hz) stand for the locations of vibration peaks.",
        "texts": [
            " In Figure 4, there exists many vibration peaks located in the speed region of 1000 r/min to 100,000 r/min as marked by symbols nxvc (4400 r/min), nyvc (6300 r/min), nxb (10,500 r/min), nyb (10,500 r/min), nxc (22,100 r/min), nyc (31,500 rpm), and nxc2 (71,400 r/min), etc. To make certain of the forming reasons of vibration peaks in Figure 4, the cascade plots respectively in the horizontal and vertical directions are obtained as shown in Figure 5. For the enormous result data, the speed interval in Figure 5 is set to be 1000 r/min, which would certainly miss many short-lived details. So, the more detailed cascade plots are shown in Figure 6 at 30,000 r/min. Combining Figures 4 and 5, it is easy to get the critical speeds of the roller bearing\u2013rotor system. Corresponding to the working frequency which is marked as f1 in Figure 5, the system gets two vibration peaks in the horizontal direction at location of Lxc (nxc\u00bc 22,100 r/min, fxc\u00bc 368.3Hz) and Lxc2 (nxc2\u00bc 71,400 r/min, fxc2\u00bc 1119Hz), and another one in the vertical direction at location of Lyc (nyc\u00bc 31,500 r/min, fyc\u00bc 524.9Hz). Thus, for the flexible rotor system with imbalance excitation, fxc and fxc2 would be the first and second critical frequencies in horizontal direction, and fyc would be the first critical frequency in vertical direction, which bring about the resonances and could be observed in Figure 4 at nxc, nyc, and nxc2",
            " From Figure 4, it is easy to find that the vibration peaks at nxvc, nyvc, nxb, and nyb all arise in the lowspeed region far away from nxc and nyc. In the work of Sinou,6 similar vibration peaks far before the critical speed could be also found in abundance, which are all due to the sub-harmonic and super-harmonic of critical frequency, and fail to be discussed with VC frequency. To explore the reason of these peaks, the detailed cascade plots respectively in horizontal and vertical directions are shown in Figure 6. Furthermore, the balanced systematic cascade plots are also obtained as shown in Figure 7 to make a comparison with an unbalanced system. Obviously, there are two vibration peaks of VC at the locations Table 1. Model parameters. Parameters Values Pitch diameter of bearing (mm) 98.5 Diameter of roller (mm) 15 Contact length of roller (mm) 14.8 Diameter of rotor (mm) 70 Diameter of disk (mm) 140 Modules (N/m) 2.0 1011 Roller number 12 BN 5.079 Length of rotor (mm) 500 Length of disk (mm) 60 Density (kg/m3) 7800 Poisson\u2019s ratio 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001088_s11044-014-9417-8-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001088_s11044-014-9417-8-Figure13-1.png",
        "caption": "Fig. 13 (Left) Humanoid upper body turning a valve. The system is described by n = 13 generalized coordinates. Generalized coordinates q1, . . . , q12 are associated with the humanoid structure and q13 is associated with the passive valve. Constraints, \u03c6(q), are defined between the humanoid hands and the valve. In this case n = 13, mC = 10, p = 3, m = 1, N = 2, and k = 12. (Right) The humanoid is commanded to turn the valve while a set of conditions, A\u03bb = d, are specified on the constraint forces. The time response of the valve angle shows linear critically damped motion to the goal. The control gains are Kx = 100 and Kv = 20. The time responses of the first six controlled constraint forces (in the global reference frame) are plotted as well",
        "texts": [
            " In the case that k > m (task actuated) and we choose to control only the task, we have \u03c4 + T \u03bb \u2212 U\u0302T c \u03c4N = \u0302 T JT ( \u0302cf + \u03bc\u0302c + p\u0302c ) + T (\u0302\u03b1 + \u03c1\u0302), (72) complemented by the following conditions on the constraints forces; A(q, q\u0307)\u03bb = d(q, q\u0307), (73) where A(q, q\u0307) \u2208R (k\u2212m)\u00d7mC and d(q, q\u0307) \u2208R k\u2212m. The passivity constraints are Sp \u03c4 = 0. (74) This can be expressed as the following system of n + mC + N equations: \u239b \u239d 1 T \u2212U\u0302T c 0 A 0 Sp 0 0 \u239e \u23a0 \u239b \u239d \u03c4 \u03bb \u03c4N \u239e \u23a0 = \u239b \u239d s(q, q\u0307) d 0 \u239e \u23a0 . (75) As an illustrative example we consider the humanoid torso depicted in Fig. 13 turning a valve. The system is described by n = 13 generalized coordinates. The constraint equations associated with the loop closure are \u03c6(q) = \u239b \u239c \u239c\u239c \u239c\u239c \u239c \u239d rrh \u2212 rvrh rlh \u2212 rvlh \u03b1rh \u2212 \u03b1vrh \u03b2rh \u2212 \u03b2vrh + \u03c0/2 \u03b1lh \u2212 \u03b1vlh \u03b2lh \u2212 \u03b2vlh \u2212 \u03c0/2 \u239e \u239f \u239f\u239f \u239f\u239f \u239f \u23a0 , (76) where rrh and rlh are the contact locations on the left and right hand, respectively, and rvrh and rvlh are the corresponding contact locations on the valve. The terms \u03b1 and \u03b2 denote the xy Euler angles in an xyz sequence describing the orientation of the hands and the valve. We note that mC = 10 and p = 3. Considering all joints except the valve joint to be actuated, we have k = 12 and Sp = ( 0 0 0 0 0 0 0 0 0 0 0 0 1 ) . (77) We will define the task to control the valve angle (see Fig. 13). So, m = 1, N = 2, and x q13 = \u03b8. (78) We will specify the reference value as xd = 0.18 rad. (79) The null space torque will be specified to be zero (\u03c4N = 0). Additionally, we wish to specify the first nine constraint forces. Thus, A = \u239b \u239c\u239c \u239c\u239c \u239c\u239c \u239c\u239c \u239c\u239c \u239c\u239c \u239d 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 1 0 \u239e \u239f\u239f \u239f\u239f \u239f\u239f \u239f\u239f \u239f\u239f \u239f\u239f \u23a0 . (80) We will specify the reference value of the constraint forces as d = \u239b \u239c\u239c \u239c\u239c \u239d Qy(q13)(\u22125,1,\u22121)T Qy(q13)(5,1,1)T 0 0 0 \u239e \u239f\u239f \u239f\u239f \u23a0 . (81) That is, the desired constraint forces are specified in the local valve reference frame and transformed into the global frame. The linear (PD) control law of (32) is used as the input of the decoupled system. The gains are chosen so as to achieve critically damped behavior of the task motion. Equation (71) is solved to compute the control torque. A small dissipative term has been added to the null space. Figure 13 shows simulation plots for the system under goal position commands on the task coordinate, x. The time response of the valve angle shows linear critically damped motion to the target. The time response of the first six controlled constraint forces (in the global reference frame) are plotted. A general methodology for handling constraints will now be described that aggregates the persistent (bilateral) constraints associated with the robot mechanism and the transient (unilateral) constraints associated the interaction of the robot with the environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000495_j.protcy.2015.07.023-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000495_j.protcy.2015.07.023-Figure3-1.png",
        "caption": "Fig. 3. 4D vector diagram of the kick force displayed on a virtual model of a soccer boot (the rainbow-coloured insert refers to the time scale).",
        "texts": [
            " The row number of each vector is converted to an AutoCAD colour in MS Excel with the equation CEILING(ROW()*24/L,1)*10 (1) where L is the last row number of the data array (and the first row number = 1). The Excel data are converted to a script file in MS Word (.scr) and the file is uploaded in AutoCAD, and added to a virtual soccer boot model. 3. Results The movement of the COP showed a consistent path of the curve in both kicks (Fig. 2), moving backward first, and reversing its direction at the peak force. Surprisingly, the COP moves only within 0.01 m. The force vector diagrams are shown in figure. 3. The COP is located on the inner side of the instep, the typical contact area of curve kicks. In spite of the low costs of the sensing material, the smart soccer boot delivers highly accurate and repeatable data. The smart soccer boot is useful for counting the number of kicks, assessing the magnitude of the kick force, displaying the COP on the boot and correlating the COP and force to the kicking accuracy. Therefore, it is believed that the system can be used commercially in the future in combination with an audible/ tactile / visual biofeedback system and smart phone application to reflect and enhance kicking performance in athletes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000505_b978-0-12-417049-0.00012-2-Figure12.16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000505_b978-0-12-417049-0.00012-2-Figure12.16-1.png",
        "caption": "Figure 12.16 Desired trajectory of the WMR.",
        "texts": [
            "90) for \u00bd\u03bc;\u03c9 T 5 \u00bd_z;\u03c9 T we get: _z \u03c9 5 cos \u03c6 sin \u03c6 2\u00f0sin \u03c6\u00de=z \u00f0cos \u03c6\u00de=z w1 w2 which is the desired dynamic state feedback linearizing and decoupling controller: _z5w1 cos \u03c61w2 sin \u03c6 \u00f012:93a\u00de v5 z \u00f012:93b\u00de \u03c95 \u00f0w2 cos \u03c62w1 sin \u03c6\u00de=z \u00f012:93c\u00de with: \u20acy1 5w1; \u20acy2 5w2 \u00f012:93d\u00de Numerical Results\u2014We consider the following values of the system parameters and initial conditions: Initial WMR position: xQ\u00f00\u00de5 1:4 m; yQ\u00f00\u00de5 1:3 m;\u03c6\u00f00\u00de5 45 Sonar position in Qxryr : xr;1 5 0:5 m; yr;1 5 0:5 m; \u03c6r;1 5 0 Plane position: p1n 5 7:0 m; \u03b81n 5 45 Disturbance/noise: Q5 diag\u00bd0:1; 0:1; 0:1 ; R5 diag\u00bd1023; 1023; 1023; 1023 The desired trajectory starts at y1;d\u00f00\u00de5 xQ;d\u00f00\u00de5 1:5 m, y2;d\u00f00\u00de5 xQ;d\u00f00\u00de5 1:5 m and is a straight line that forms an angle of 45 with the world coordinate Ox axis as shown in Figure 12.16. The new feedback controller \u00bdw1\u00f0t\u00de;w2\u00f0t\u00de T is designed, as usual, using the linear PD algorithm: w1 5 \u20acy1;d 1Kp1\u00f0y1;d 2 y1\u00de1Kd1\u00f0 _y1;d 2 _y1\u00de w2 5 \u20acy2;d 1Kp2\u00f0y2;d 2 y2\u00de1Kd2\u00f0 _y2;d 2 _y2\u00de Figure 12.17A shows the trajectory obtained using odometric and sonar measurements and the desired trajectory. Figure 12.17B shows the desired orientation \u03c6d, and the real orientation \u03c6 obtained using the EKF fusion. As we see, the performance of the EKF fusion is very satisfactory. Example 12.6 It is desired to give a solution to the problem of estimating the leader\u2019s velocity vl in a leader follower vision-based control system (Figure 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure13.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure13.2-1.png",
        "caption": "Fig. 13.2 Modes of operation; (i) flow mode, (ii) shear mode, (iii) squeeze mode",
        "texts": [
            " By addition of additives and surfactants synthesized MR fluid undergoes minimal properties degradation over the time. Concluding the literature study, considering the application, mode of operation and operating conditions MR fluid should be specially designed. As every application has its own shear rate, yield stress, temperature stability and acceptable sedimentation rate. The application range of Magneto-rheological fluid being wide engineering applications, the working principle or modes of operations of MR fluid is categorized into three classes (Fang et al. 2009), as shown in Fig. 13.2 \u201cflow mode\u201d is represented where the flow of the fluid is along the constrained surface also normal to that of externally applied magnetic field as demonstrated in Fig. 13.2(i). whereas the \u201cshear mode\u201d is categorized with one of the surface moving with a velocity in one direction along the flow with other surface fixed and the flow is normal to applied magnetic field as demonstrated in Fig. 13.2(ii). Kim et al. (Kim et al. 2012) also the \u201csqueeze mode\u201d quite same as shearmodewith a difference of the surfacemovingwith velocity is along the normal to other surface and also parallel to externally applied magnetic field as demonstrated in Fig. 13.2(iii). The Magneto-Rheological fluid offers solution wide range of engineering application. The development of MR fluid technology is evident in numerous specializations, extending from structural building and automotive to biomedical designing applications. Innumerable Experimental has been reported which distinguish the 13 Magneto Rheological Fluid Based Smart Automobile Brake \u2026 243 advantages or utilizing MR gadgets in different fields (Mazlan et al. 2008). This allaround reported achievement of MR fluids keeps on persuading present and future uses of MR fluid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure4.21-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure4.21-1.png",
        "caption": "FIGURE 4.21 Capacitor with an electrode rotated about an arbitrary point.",
        "texts": [
            " The data file used is shown in Table 4.4. Figure 4.20 shows the result of this rotation. 84 Examples Using the Method of Moments electrode), it is necessary only to define that point as the center of the coordinate system. TABLE 4.4 Data File for Capacitor Shown in Figure 4.19 0.01 0 0 25 200 0 \u22120.5 45 0 0 25 200 0.1 0.5 0 260 240 220 200 180 160C ( pf d) 140 120 100 80 0 20 40 60 80 \u03b8 (\u00b0) 100 120 140 160 180 FIGURE 4.20 Capacitance as a function of rotation angle \u03b8 for capacitor shown in Figure 4.19. 4.8 Varying the Geometry 85 Figure 4.21 shows an example of this situation; with respect to the origin, the lower left corner of the electrodes is (\u22120.25,\u20130.25). The data file is shown in Table 4.5, and the capacitance as a function of \u03b8 is shown in Figure 4.22. Figure 4.22 shows a curious capacitance behavior\u2014while it is expected that the capacitance will fall drastically as \u03b8 is increased from 0, the capacitance variation TABLE 4.5 Data File for Capacitor Shown in Figure 4.22 0.01 \u22120.25 \u22120.25 20 160 0 \u22120.5 45 \u22120.25 \u22120.25 20 160 0.1 0.5 0 180 160 140 C 120 100 80 60 0 20 40 60 80 Theta 100 120 140 160 180 FIGURE 4.22 Capacitance as a function of rotation angle \u03b8 for capacitor shown in Figure 4.21. 86 Examples Using the Method of Moments near \u03b8 = 180 degrees might be unexpected. Examination of Figure 4.22, however, shows that this behavior is perfectly reasonable because of the electrodes\u2019 overlay area as \u03b8 is varied. The capacitor shown in Figures 4.21 and 4.22 is clearly a poor choice as a variable capacitor for tuning a radio or as a rotation angle encoder; monotonic capacitance versus rotation angle behavior is mandatory in these applications. This type of information is a very valuable result of simulation; we usually can forgive an absolute capacitance error of a few percent as long as this error repeats in a production environment, but surprises such as the behavior shown Figure 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002632_tencon.2016.7848445-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002632_tencon.2016.7848445-Figure4-1.png",
        "caption": "Fig. 4: Stereolithographic parts",
        "texts": [
            " The whole process is the repetition of the above step till the whole part being built. During the building of parts, design should be made carefully to avoid internal and overhanging structures. After printing, the platform is raised and the printed parts are cured. The process of curing is done is oven like machine that makes the resin harder and layers stronger[8]. The limitation of this process is the stability of the products over time, as the layers may peel of to make the parts more brittle. Example of parts which are manufactured using SL are shown in Fig. 4. Selective laser sintering(SLS)[9] or laser melting is a process which very similar to the previous process only that the material used are in powder form(SLS process is shown in Fig. 5). The powder can be either metal or polymer depending on the application. SL machine has a high power laser on the top of the machine that will melt the powder to fuse layer over layer. As the melting point of the metal powder is very high, power of the lasers is considerably larger than the SL process. The SLS process takes the STL code and builds the parts inside a powder bed with the moveable platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001756_978-3-319-24502-7_10-Figure10.29-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001756_978-3-319-24502-7_10-Figure10.29-1.png",
        "caption": "Fig. 10.29 Left Two pushers (p > 0) on a not too diverging course attract each other hydrodynamically, and reorient each other into a parallel side-by-side motion. Right Two pullers (p < 0) on a diverging course reorient each other towards an antiparallel configuration, swimming subsequently away from each other in anti-parallel direction (see also [108])",
        "texts": [
            " A freely advecting sphere for which C \u00bc 1 rotates thus with half the vorticity of the incident flow uinc, taken at the sphere center. An elongated body has an additional angular velocity part due to the shear strain part of the incident flow. This additional angular velocity part is oriented perpendicular to the long body axis unit vector e1 of swimmer 1. Substitution of uD(R12;e2) for the incident flow into Eq. (10.124) reveals that two nearby pushers on a converging course reorient each other hydrodynamically into a parallel side-by-side configuration. As depicted in the left part of Fig. 10.29, if two pushers located in the x \u2013 z plane are separated by the distance h = R12, and symmetrically oriented with inclination angles h relative to the z axis, their reorientation into a parallel configuration takes place with the angular velocities Xy;12 ph=\u00f0gh3\u00de [108]. In contrast, and owing to the opposite flow fields, two pullers on a diverging course align each other in an antiparallel configuration, swimming subsequently away from each other. This is illustrated in the right part of Fig. 10.29. It should be recalled that the analysis presented here is based on the leading-order singularity flow solutions. It applies in principle to inter-swimmer distances h large compared to the elongational swimmer size L only, although it is often found to be quite accurate even for distances comparable to L [109]. For two closely moving swimmers, the details of their shapes and propulsion mechanisms play a role. This requires then a refined hydrodynamic modeling and more elaborate methods to determine the swimmer dynamics",
            " The image flow contribution is required to satisfy the surface BC (recall Sect. 10.5). In Fig. 10.30, this situation is illustrated for the simplest case of a pusher swimming above a free surface. The only BC here is the fluid-impermeability of the surface which can be fulfilled by considering the swimmer in the half-space z > 0 to move along with its mirror image in the fluid extended to the lower half-space z < 0. This is akin to the symmetric side-by-side motion of two swimmers in bulk fluid discussed earlier in relation to Fig. 10.29. Using again the translational Fax\u00e9n theorem for a point-like freely advected particle, the vertical velocity component induced on the swimmer at R0 = (0,0,z0) with z0 = h > 0 is Vz\u00f0h; h\u00de \u00bc uD;z\u00f0R0 R 0; e im\u00de \u00bc p 32pgh2 1 3 sin2 h ; \u00f010:125\u00de where h is the tilt angle of the swimmer with respect to the surface, so that sin h \u00bc e ez \u00bc eim ez. Here, R 0 and eim are the position and orientation vectors of the mirror dipole, respectively. Provided the tilt angle is not too large so that h\\ arcsin\u00f01= ffiffiffi 3 p \u00de, the dipole is attracted by the surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000840_00423114.2015.1023319-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000840_00423114.2015.1023319-Figure7-1.png",
        "caption": "Figure 7. Multibody model of the 18\u2013100 bogie.",
        "texts": [
            " The scheme of the load application should provide a realistic stress\u2013strain state in all elements of the model. The load is applied as a system of concentrated vertical forces as shown in Figure 6(b). Two torques are resulted D ow nl oa de d by [ G az i U ni ve rs ity ] at 1 1: 00 3 0 A pr il 20 15 Vehicle System Dynamics 863 from the difference between the corresponding vertical forces. The values of the loads come from the dynamics simulation. 5.2. Multibody model of the 18\u2013100 bogie for dynamics simulation The multibody model of the 18\u2013100 bogie (Figure 7) was developed in the Universal Mechanism (UM) software package and described in detail.[3] UM is general-purpose software for multibody system dynamics analysis often used for the simulation of railway vehicle dynamics. The model includes four friction wedges as rigid bodies with six degrees of freedom; the model\u2019s total number of degrees of freedom is 54. The bogie has rigid contacts between the side frames and the axle boxes, between the car body and the bolster in the centre plate and side bearings (with a clearance), and between the wedges and the bolster or the side frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure12-1.png",
        "caption": "Fig. 12 Compound twocarrier planetary gear trains with one compound and four external shafts. a One-engine driving of two opposite rotating aircraft propellers (separating train). b Oneengine driving of asymmetric differential of a heavy vehicle (separating train). c Twoengine driving of aircraft flap (collecting train)",
        "texts": [
            " In the gear trains with internal power circuits it is shown that with one and the same structural scheme, the kinematic schemes may differ (Fig. 7). Figures 10 and 11 show two examples of the compound two-carrier planetary gear trains with two compound shafts and four external shafts. The brakes used in these gear trains can be located on different shafts\u2014single or compound. The Fig. 10 shows a reverse gear train, carrying forward and backward. The Fig. 11 shows a change-gear (gearbox), carrying out two gear-ratio steps (speeds). Figure 12 shows three cases of using two-carrier compound planetary gear trains with a one compound shaft, but with four external single (i.e. not-compound) shafts. Figure 13 shows the structural schemes of three-shaft three-carrier and fourcarrier compound gear trains. The first are used as reducers or multipliers, i.e. with F = 1 degree of freedom, while the second\u2014as change-gears. Figure 14 shows three-carrier compound gear train that works as a multiplier in the powerful wind turbine (Giger and Arnaudov 2011)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.22-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.22-1.png",
        "caption": "Figure 2.22 Earlobe electrodes (A) with soldered lead wire and (B) with lead wire connector (the plastic clip has a top hole coinciding to the electrode central hole to apply the electrolyte gel).",
        "texts": [
            " As a result, these electrodes are homogeneous across the thickness, do not bend, have high mechanical strength and do not require rechloriding. They have wide contact area and present as further characteristics, very low offset voltage, polarization, rate of drift and noise level and are less susceptible to artifacts than the conventional electrodes (Tallgren, Vanhatalo, Kailaa, & Voipio, 2005). Figure 2.21 shows some examples of sintered EEG electrodes. Figure 2.21A shows sintered fixed wired electrode without housing and Figure 2.22B shows the electrode with plastic cover. Sintered discs are manufactured with and without a central hole (Figure 2.21C), in different disc diameters (4 12 mm) to suit to infant and adult use; they are coupled to the scalp skin with electrolytic gel and are fixed using collodion like the conventional electrodes. They are widely used with caps for brain mapping and long-term brain activity monitoring. The noninvasive EEG recording routines include some types of nonscalp electrodes, like the earlobe clips, used as reference, and nasopharyngeal electrodes, used in investigation of the origin from specific epilepsy activities. The earlobe electrode has two EEG sensors made of gold, silver, or Ag AgCl, mounted on a clip type plastic holder showed in Figure 2.22A. The electrical connection is made by soldering the lead wire to the metal sensor (Figure 2.22A) or by a conector (Figure 2.22B). The metallic sensor is manufactured with different disc and hole diameters and cable lengths (same as EEG electodes) to suit to infant and adult patients. The discs must be made from the same material of the scalp electrodes to avoid measurement artifacts caused by dissimilar electrodes. Nasopharyngeal electrodes are rigid and have a format similar to the letter z with the first and central (10 15 cm) parts made of insulated lead wire (usually silver) and the third part (tip, 2 3 cm) coated with gold ending in a ball (active region) with 2 5 mm diameter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000072_s12206-019-1010-x-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000072_s12206-019-1010-x-Figure6-1.png",
        "caption": "Fig. 6. Experiment on mechanical structure connected by bolts.",
        "texts": [
            " The material densities for elements A and B are 1r and 2 ,r respectively, and the average virtual material density of the fixed joint interface is then 1 2 1 1 2 2 j 1 2 1 2 1 j 1 2 j 2 1 1 2 2 j 1 j 2 1 2 m m \u03c1 v \u03c1 v\u03c1 v v v v \u03c1 A h \u03c1 A h \u03c1 h \u03c1 h A h A h h h + + = = + + + + = = + + (22) where jr is the virtual material density of the fixed joint interface; 1m and 2m are the masses of the surface layers for elements A and B, which are the components of the fixed joint interface; 1V and 2V are the volumes of the surface layers for elements A and B, which are the components of the fixed joint interface; and jA is the macroscopic area of the fixed joint interface [1]. In general, the metallic material properties, manufacturing process, and other conditions for the surfaces of two elements forming a fixed joint interface are identical; thus, h1 = h2 is assumed, and Eq. (22) can be rewritten as .1 2 j \u03c1 \u03c1\u03c1 2 + = (23) Considering the characteristics of fixed joint interface, a mechanical structure connected by bolts was analyzed. The experimental equipment schematic is shown in Fig. 6(a). The fixed joint interface consisted of the contact part from experimental block 1 and block 2. Both experimental block 1 and block 2 are made of No. 45 steel with an elastic modulus E = 2.06\u00d7105 MPa and a Poisson\u2019s ratio of 0.26.m = The fixed joint interface of experimental block 1 and block 2 was processed by abrasive machining with a roughness degree of 0.8 m,m and was dry contact and oil free. Experimental block 2 was fixed and immovable. In the experiment, the torque wrench rotated the bolt through a single-direction force sensor, and applied loading to the experimental subject through the gasket. The deformation of each test point was measured by a differential inductance sensor. The testing points were located at the top surfaces of experimental block 1 and block 2. The experiments used a GB/T93-12-type spring gasket. The experimental bolts are GB/T5781-M12\u00d750-type six-angle-head, full-thread bolts. The clearance between the six-angle-head threaded bolt and test block 1 was clearance fit. Fig. 6(b) shows the mechanical characteristics analysis and test field diagram of the bolt connection. Fig. 6(c) shows the related dimension diagram of experimental block 1 and block 2. Fig. 6(d) shows an overhead view; a, b, and c mark various testing points. Fig. 7(a) shows the experimental calculation model of the mechanical structure by bolted connection. The test block 2 and the bolt portion that is connected to test block 2 are simplified into a component A. The fixed joint interface consists of the contact portion of experimental block 1, and block 2 was defined as element B. The remaining portion of experimental block 1 was defined as element C. The top surface of element C was subject to a uniformly distributed annular load q"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001435_iecon.2013.6699639-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001435_iecon.2013.6699639-Figure9-1.png",
        "caption": "Fig. 9. For the case of one broken bar: a) Magnetic flux density; b) Current density distribution.",
        "texts": [
            " The magnetic flux density and current density distributions for a healthy rotor are shown in Fig. 7. The radial and tangential forces are shown in Fig. 8. It can be seen from these figures that, for the case of healthy rotor, the radial forces are almost symmetric and cancel each other. On the other hand, the tangential forces, which are responsible for generating the motor torque, are symmetrical under the machine poles. The magnetic flux density and current density distributions for the case of one broken bar are shown in Fig. 9. Distributions of the tangential and radial forces corresponding to the same instant are shown in Fig. 10. It can be observed from Fig. 10(a) that the radial forces are asymmetrical in this case, leading to a net radial force affecting the rotor shaft and, eventually, the bearings. Moreover, the tangential forces are no longer symmetric under the machine poles as shown in Fig. 10(b). The negative tangential forces cause a decrease in the average torque compared to the healthy case. Moreover, distributions for the magnetic flux density and the current density, for the case of two adjacent broken bars, are shown in Fig",
            " It can be observed from Fig. 12(a) that the asymmetry in the radial forces has increased in comparison to the case of one broken bar. Such increase in the asymmetry would negatively affect the bearings life time. Increased asymmetry of the tangential forces, as illustrated by Fig. 12(b), would result in increased torque and speed variations. The negative tangential forces cause a decrease in the average torque compared to the healthy case. It can be seen from the magnetic flux density distributions, shown in Fig. 7(a), Fig. 9(a) and Fig. 11(a) that the distribution in the faulty cases are asymmetric compared to the healthy case. This paper presents an efficient model to analyze broken bar fault conditions for induction motors using a FEM approach coupled with an ABC transient model. The pattern of the asymmetry in bar currents resulted from various broken bar faults can be deduced and analyzed. Effects on the motor torque and speed can be observed. Radial and tangential force asymmetries resulting from broken bars conditions can be assessed using the proposed approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002008_jahs.60.022012-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002008_jahs.60.022012-Figure2-1.png",
        "caption": "Fig. 2. Raptor 50 V2 rotor and stabilization system.",
        "texts": [
            " The stabilizer bar is an essential part of small-scale helicopters, which have higher sensitivity to control inputs than their full-scale counterparts. Without some form of stability augmentation, it is almost impossible for human pilots to control such vehicles. Various methods have been employed to increase the stability of these model-scale helicopters. The stabilizer bar is one of the most widely used devices for the stability augmentation of model-scale helicopters. The system consists of flybars and paddles placed at 90o to the main rotor blades as shown in Fig. 2. The bars are attached to the rotor shaft above the main rotor through an unrestrained teetering hinge and are coupled to the main rotor through a Bell mixer. The stabilizer bar acts as a lagged rate feedback in the pitch and roll axes. This reduces the bandwidth and control sensitivity to cyclic lateral and longitudinal inputs. Using stabilizer bars, cyclic commands do not go directly to the blades. Instead, the cyclic commands are applied to the flybar whose flapping motion determines the blade pitch angles (Refs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002533_iros.2016.7759198-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002533_iros.2016.7759198-Figure1-1.png",
        "caption": "Fig. 1. Tracked vehicle with sub-tracks \u201dQuince\u201d and its sub-track attached with caster wheels on the side surface",
        "texts": [
            " However, daily inspection is arduous owing to the limitations of human resources and cost. In addition, human inspectors have a high risk of accidents during their daily inspections of production lines. Thus, it is necessary to develop an inspection robot that inspects such environments instead of a human. Autonomous inspection robots can solve the problems arising from a lack of human resources and safety issues. We have developed an autonomous inspection system for industrial plants using the tracked vehicle shown in Fig.1(a). The tracked vehicle was used for inspections because it can drive on uneven floors in the production lines. The tracked vehicle autonomously runs in passages and monitors the facility that contains the product lines. Autonomous inspection in an industrial plant is a challenging task for robots because footholds are uneven and the passages for inspection are narrow. Using a tracked vehicle mechanism does not ensure accurate position control because tracks always slip while navigation in principle of the mechanism",
            " However, the tracked vehicle cannot completely avoid collisions when it drives near walls because of slips and sensing limits. We propose a motion control method based on a dynamic model of a tracked vehicle colliding with a wall. Side boards are attached to the handrails of the inspection passages to prevent falls. They are also attached to the walls to prevent damage. We use these boards and walls to autonomously move in the direction of the target. The proposed method has three features. First, passive wheels are attached to the side surfaces of the tracked vehicle, as shown in Fig.1(b). This helps to avoid damage to the wall and to move the vehicle in the direction of the target. Second, the tracked vehicle is controlled based on a dynamic model that uses driving force and collision force. Third, this method can be applied to robots that do not have accurate force sensors attached to the active flippers, as the method only requires the simple information of whether the robot is colliding or not. This information can be obtained using basic internal sensors such as IMU sensors or current sensors of the driving motors",
            " We propose a control method for a tracked vehicle in narrow space. As shown in Fig.2, the tracked vehicle autonomously moves along target paths during navigation. The vehicle also moves along the target paths whenever it makes contact with walls. Although a position error remains, the tracked vehicle can follow the target paths. This is an important function in narrow passages. To reduce damage to walls caused by collisions with the tracked vehicle, caster wheels are inserted on the side of the body, as shown in Fig.1(b). The use of caster wheels has the following advantages: the caster wheels can reduce the friction force between the tracked vehicle and the wall, can move the robot over small bumps along the wall, and can passively adjust the pose to synchronize with the direction of the friction force. In addition, the soft wheels can reduce any damage to the walls. The tracked vehicle, which makes contact with the wall by using the caster wheels, can now smoothly follow along the wall. However, the contact force prevents the tracked vehicle from leaving the wall because the contact force generates a rotational moment in the opposite direction of the moving direction when the tracked vehicle leaves the wall"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001365_s11041-014-9680-6-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001365_s11041-014-9680-6-Figure8-1.png",
        "caption": "Fig. 8. Distribution of temperature over an aluminum ingot 100 mm long and 60 mm in diameter (the beam power is 7.5 kW, the diameter is 20 mm, the growth rate is 3 mm min): h ) height of the ingot; l ) distance from the center of the ingot; 1 ) contour of the molten zone; 2 ) plane of the melting temperature (932 K).",
        "texts": [
            " The heat flows computed for three growth rates are given in Table 4. It can be seen that the main loss in the energy is caused by the heat flow through the ingot crucible wall interface (G2). A considerable part of the energy loss (about 17 \u2013 25%) corresponds to the flow through the bottom of the crucible (G4). An advantage of the model used is the possibility of three-dimensional graphical visualization of the distribution of temperature in a metallic ingot and of simple interpretation of these results. For example, Fig. 8 presents the distribution of temperature in an aluminum ingot due to drip melting. The cross section of the three-dimensional distribution of the temperature and the plane of the melting temperature give the contour of the molten zone (Fig. 8). The results of the computations can be presented in another way (Fig. 9), i.e., in the form of the curves of temperature distribution at three heights of the aluminum ingot under the same conditions as in Fig. 8. A model for describing heat transfer in electron-beam melting of aluminum by the method of drip melting and of tantalum by the method of disk melting has been suggested and checked. Data have been obtained on the geometry of the solidification front, on the thickness and width of the molten zone, on the heat losses through the interfaces of the ingot at various values of the power, of the radius of the electron beam and of the rate of growth of the ingot. The computed and experimentally determined geometries of the molten zone of tantalum agree well"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000870_infcom.2013.6567020-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000870_infcom.2013.6567020-Figure7-1.png",
        "caption": "Fig. 7. P(, Qi are in level-i; Pj+i, Qi+i are in level-(\u00a3 + 1). (a) The component P; is connected with P; + i and Q,+\\; (b) The component P; corresponds to component Pj + i only; (c) The component Pj+i is connected with P; and Qj,",
        "texts": [
            ")\u2022 if there exists a node p in the component Pt that has a neighbor // in the component Pl+\\. Then the nodes p and p' will notify the dominating landmarks DM (Pi) and DM(Pi+l) in the components Pi and P;+i respectively of this connectivity. According to Morse theory, there only exist three cases: (1) Pi is connected with Pi+} and another component Qi+\\ in level(i + 1); (2) Pi corresponds to Pi+\\ only; (3) P,+i is connected with Pi and another component Q, in level-?\", which is notified to DM(Pi) and DM{Pi+]) respectively. The three cases are illustrated in Fig. 7. The dominating landmarks will notify the nodes in the component Pi+\\ if P,+ ] only corresponds to the component P,. If so, the nodes in Pj+i are assigned the same region ID as the nodes in component Pi, otherwise, the nodes in component Pj+i are assigned a new region ID. After this process, every node is notified with its region ID. The result of constructing the Reeb graph is given in Fig. 5(c), the Reeb graph regions are distinguished by colors. With the Morse function and Reeb graph constructed, all the loop-end regions are notified in a straight forward manner: if the dominating landmark DM(Pi+i) in Pi+\\ is notified that P,_I_I is connected with P, and Q, in level-/ (as shown in Fig. 7(c)), then component Pt+I and all the components in the same region are notified to be in a loop-end region. Then to extract the maximum cut set Cmax. each loop-end region /*\u201e performs a bisection operation to extract a cut Q as follows, Suppose a loop-end region Plu consists of a set of components {Ph} where H\\ < h < H2, then the bisection operation in / ' n is initiated from Pa, to P//,. The nodes in PH, reset its region ID to its parent's region ID. Since P//, is the first component of Pm, it is the neighbor of two components Pi and Qi (Fig. 7(c)), and the component p and Qi belong to region In and Ip respectively (as described in Section II-C). The nodes in P// l change their region ID to the region ID of Ia or Is respectively. Therefore, Pa, is bisected and assigned to region Iti or 1$. This process is carried out from PH^ to Ptf.2- Consequently, loop-end region 1^ is bisected and two newly merged region I'a and I'a (as described in Section II-C) are generated, as shown in Fig. 5(d). Next, the cut C; is obtained by disconnecting I'a and Ii"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002808_0959651819899267-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002808_0959651819899267-Figure6-1.png",
        "caption": "Figure 6. Arrangement for estimation of thrust and drag forces, and nonlinear cable stiffness.",
        "texts": [
            " The range of the input voltage signal to the motors given in Simulink real-time workshop is 22.5 to +2.5 V. The corresponding range of actual terminal voltages of the main motor is 217.93 to +17.58 V, and that of the tail motor is 213.12 to 12.71 V. The thrust forces and drag forces acting on the rotors at various speeds of the rotors are obtained through an experimental parameter estimation procedure. The thrust and drag forces at various positive and negative voltage inputs or rotor speeds are measured using an electronic balance, as shown in Figure 6. An initial weight is applied and the reduction in weight due to string tension is measured. For reverse rotation of main rotor, a pulley arrangement is used. Similarly, pulley arrangement is used to convert horizontal force due to tail rotor actuation to vertical force applied to the electronic balance. The measured thrust forces for main and tail rotor at various speeds are plotted in Figures 7 and 8, respectively. Least square method is adopted to obtain the thrust coefficients using equations (6) and (7)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003526_icuas48674.2020.9213837-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003526_icuas48674.2020.9213837-Figure5-1.png",
        "caption": "Fig. 5: Representation of \u03b5-ratio between any two paths. Figure also shows that path \u03a61(\u00b7) is \u03b4 -close to path \u03a62(\u00b7)",
        "texts": [
            " Intuitively speaking, if we can put an empirical bound on the value of \u03b5-ratio, then we can put a bound on how much a path is likely to diverge after it is repaired. This bound is a function of how far the endpoint of the surrogate is from the actual boundary condition, i.e. a solution is less likely to deviate from the surrogate if the boundary condition is almost satisfied. If the repaired path is \u03b4 -close to the surrogate solution, we can guarantee that the repaired path does not collide with any obstacles. Consider \u03a62(\u00b7) in Figure 5 to be the selected surrogate solution and \u03a61(\u00b7) be the corrected one. As the value of b is known, we may choose a \u03b4 inflation factor for the obstacles such that \u03b4 > \u03b5b. For such a choice of \u03b4 , any repaired path will not collide with obstacles, and a collision check on the surrogate is equivalent to a collision check for the repaired path. In practice, getting a strict bound on the \u03b5-ratio might be difficult, so we approximate the value by running multiple simulations and recording the value. In order to find the value, we ran the BVP solver in the same manner as for calculating the path length changes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002050_s10999-016-9338-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002050_s10999-016-9338-1-Figure1-1.png",
        "caption": "Fig. 1 Geometric model for normal and side contact of two asperities on rough surfaces: a geometric model of asperity normal contact and b geometric model of asperity side contact",
        "texts": [
            " The normal contact occurs at the peaks of two asperities if friction between the contact surfaces is excluded. However, the side contact occurs at a point, which deviates from the peak when friction is included. Obviously, normal contact of the asperities is an ideal state for rough surfaces, while side contact actually occurs at the interface of a mechanical higher pair. Assuming that asperities are paraboloids, the geometric models for the normal and side contact of two asperities are shown in Fig. 1a, b, respectively. In Fig. 1, R1 and R2 are the curvature radius of the two asperities, d0 the distance between the rotation axis of two asperities, and y0 is the distance between the peak of two asperities. In Fig. 1b, n and s are the normal and tangent vector on the XOY plane respectively, and n0 and s0 are the normal and tangent vector in contact point O3 respectively. F is the normal contact load in contact point, and Fn and Fs are the vertical and horizontal component of F respectively. Where l = cot h = y0/d0, and l is the average friction coefficient of the contact surfaces. The shape equations of the two asperities are y1 x\u00f0 \u00de \u00bc x2 2R1 \u00f01\u00de y2 x\u00f0 \u00de \u00bc x\u00fe d0\u00f0 \u00de2 2R2 y0 \u00f02\u00de In Eqs. (1)\u2013(2), R1 and R2 can be calculated from (Johnson and Johnson 1987)",
            " (9), the contact load of the elastic region is FTe \u00bc Z ra rp 2prr r\u00f0 \u00dedr \u00bc 4Et 3R r2 a r2 p 3 2 \u00f013\u00de The total contact load of the elastic and plastic contact regions is FT \u00bc FTe \u00fe FTp \u00bc 2Et 3R 2r2 a \u00fe r2 p r2 a r2 p 1 2 \u00f014\u00de Under the action of the normal contact load F, the plastic deformation occurs within the contact region of the radius rp, while the elastic deformation is still generated in other region. Using Eqs. (8)\u2013(9), the elastic\u2013plastic normal contact stress in normal contact is given by r1 \u00bc 3F 2pr2 a 1 r2 p r2 a !1 2 \u00f015\u00de 3.2 Elastic\u2013plastic asperity contact model with friction When friction between two rough surfaces is included, the side contact occurs on the interface of two smooth asperities, as shown in Fig. 1b. Assuming that the contact region, whose distance is y2 from the contact point O3, is composed of the plastic (inner ellipse) and elastic region (outside ellipse annulus) (see Fig. 3). Similarly, y2 can be calculated from Eq. (11). The equivalent curvature radius of the two asper- ities at the contact point in XOY coordinate plane is R0 \u00bc R 1 \u00fe d2 0 R1 \u00fe R2\u00f0 \u00de2 \" #3 2 \u00f016\u00de The critical contact deformation of the asperity in the normal direction n0 is dn0c \u00bc pkH 2Et 2 R0 \u00bc pkH 2Et 2 R 1 \u00fe d2 0 R1 \u00fe R2\u00f0 \u00de2 \" #3 2 \u00f017\u00de When dn0 [ dn0c in the direction n0, the contact asperities are in plastic deformation stage",
            " Assuming that asperities are paraboloids, the asperity geometric model is meshed by C3D8R element (Liu et al. 2014). It is noted that the appropriate displacement load is preloaded and increased gradually, to avoid the non-convergence of the finite element contact analysis on the asperities due to too large initial displacement. Assuming that the friction coefficient l is 0.1 at the meshing-out point of one tooth pair, the distance between the rotation axis of two asperities is d0 = 0.68 lm, and the distance between the peaks of two asperities is y0 = 0.068 lm, shown in Fig. 1. In addition, the material of the gears is 45 steel, whose Young\u2019s modulus are E1 = E2 = 206 GPa, yield strength is rs = 418 MPa, and Poisson\u2019s ratio are v1 = v2 = 0.3. In general, the elastic deformation is generated on the tooth meshing profiles in the working load range. According to the results of the document (Kogut and Etsion 2002) and the toughness of steel gear, the elastic deformation of the asperities on rough tooth profiles occurs on the contact surfaces when the asperity deformation d\\ 6dc, while the plastic deformation is only generated inside the contact bodies (asperities)",
            " Likewise, when m = 1 the asperity contact is from entirely elastic to elastic\u2013plastic and the plastic deformation begins to occur inside the asperities. Furthermore, when m = 2 larger plastic deformation is generated inside the asperities, and when m = 6 the plastic deformation is found to expand to the contact surfaces of the asperities. However, when m = 10 larger plastic deformation covers both inside the asperities and on their surfaces. 4.3 Result analysis Von Mise stresses of the asperities along the normal vectors n and n0 (seen in Fig. 1) are calculated, as plotted in Fig. 8. The extreme points of Von Mise stresses are found under various displacement loads (including 2dc, 6dc, 7dc, and 10dc), since the stress firstly increases and then decreases with the increase of the normal displacement. It is also indicated that the extreme points are located in the limit region of the asperities plastic deformation, and that the plastic deformation region expands from the inner of the asperity to its surface with the increase of the displacement load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000228_j.actaastro.2019.12.021-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000228_j.actaastro.2019.12.021-Figure3-1.png",
        "caption": "Fig. 3. Structure of the TER tracker.",
        "texts": [
            " When the SSPCI requires to turn to face the ground161 station for communication, the TER tracker keeps the elevation angle in the162 process of tracking the sun unchanged and rotates along axis Yb back to the163 direction of axis Zb, while the CMR rotates around the rotary joint at the164 lower end of the central cylinder until the central cylinder points towards the165 center of the earth, as shown in Fig. 2(b).166 7 (a) (b) 2.1. Description of three-extensible-rod tracker167 The TER tracker is designed as a simple and lightweight structure, as168 shown in Fig. 3. The TER tracker includes a base platform, a mobile plat-169 form, and three extensible rods. Both the base platform and mobile platform170 are like a wheel, which is lighter than a solid platform. Three extensible rods,171 all with the same structure, are arranged between the two platforms in an172 equilateral triangle. The lower end of each extensible rod is connected to the173 base platform by a rotary joint, and the upper end is connected to the upper174 platform by a smart compound joint that consists of a hook joint and rotary175 joint, whose axes converge at one point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000822_j.proeng.2014.12.047-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000822_j.proeng.2014.12.047-Figure3-1.png",
        "caption": "Fig. 3. Infinitesimal mode.",
        "texts": [
            " The analysed module has one infinitesimal mechanism and a corresponding self-stress state, which stiffens the module. Cable removal causes the module to collapse. The further study considers a sample plate structure constructed from reversed 3-strut Simplex modules (described above), located in such a way that the left module is always joined with the right and vice versa (Fig. 2). No additional cables were used in this structure. The plate is supported in four nodes. It has one infinitesimal mechanism (Fig. 3), and self-stress states in each module [2]. Calculations of such structures cannot be performed using geometrically linear theory because of the infinitesimal mechanism. Finite element analysis according to the second order theory allows to consider prestressing forces (in one module or several) through geometric stiffness matrix, and to obtain structural response for external loading. Fully nonlinear analysis leads to big differences within the scope of small prestressing forces, which disappear along with an increase of force values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure13-1.png",
        "caption": "Fig. 13 Cutting positions at the cradle angle /c \u00bc 0 on the bevel gear-cutting machine: (a) right flank cutting (upper position) and (b) left flank cutting (lower position)",
        "texts": [
            " In the generating method, in contrast, the cradle is rotated to allow the generating gear to roll with the work gear, meaning that the cradle angle /c gradually increases or decreases to achieve a generating movement. Here, the sliding base feed setting DB is used to control cutting depth. Tables 4 and 5 list the finishing coordinates of the five axes for the pinion and gear, respectively, each represented as a polynomial function of the cradle angle. Because only a cutter is employed, the right and left flanks are cut in the upper and lower positions (verified by SOLIDWORKS and illustrated in Fig. 13), but in contrast to the traditional machine, the CNC machine allows freeform flank modification. These coordinates can be used to program the NC data needed by the VERICUT software to confirm the correctness of machine positioning in the manufacturing stage. Figure 14 shows the cutting simulation for the pinion using VERICUT. The resulting tooth surfaces are saved as an STL (stereolithography) file and later compared with the theoretic tooth surfaces. Figure 15 shows the flank topographic deviations between the theoretic and produced tooth surfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000506_978-3-7091-1379-0_2-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000506_978-3-7091-1379-0_2-Figure4-1.png",
        "caption": "Figure 4: Distribution of the masses in a leg.",
        "texts": [
            " Calculations of velocities via screw theory method have shown the same result as the calculations via Jacobian method. The general conclusion for this section is that the proposed mechanism has no Type 2 singularities and all Type 1 and Type 3 singular points lie on the theoretical edge of the mechanism\u2019s workspace. The Lagrange-D\u2019Alembert principle was used to analyze dynamics of the mechanism. In order to simplify the calculations, we are assuming that the masses of the links (m1,m2,m3,m4) in each leg and the mass of the moving plate mP are distributed as shown in Figure 4 With this assumption the basic system of equations (in matrix form) which represents the dynamics of the mechanism can be written as follows: J   M1 +M1 Fg +M1 FI M2 +M2 Fg +M2 FI M3 +M3 Fg +M3 FI  + ( m3 +mP + m2 + 4m4 2 )  gx \u2212 ax gy \u2212 ay gz \u2212 az   = 0 (7) Here Mi is a torque in actuated joint of i\u2212th leg; M i Fg and M i FI are torques produced in i\u2212th leg by gravity force and inertia forces, respectively; ax, ay, az are accelerations of the moving plate along axes x, y, z respectively; gx, gy, gz are parts of the gravitational acceleration; J is the Jacobian matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003033_iet-epa.2019.1019-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003033_iet-epa.2019.1019-Figure3-1.png",
        "caption": "Fig. 3 Calculation results of rotor stress and deformation under 1.2 times rated speed",
        "texts": [
            "7 MPa and is around half of the yield strength of the stain steel sheet. The maximum deformation appears at the poles of the rotor, which is 0.0148 mm and much smaller than 10% of the air gap length, ensuring the safe operation of the motor. Furthermore, it has been pointed out in the related standard IEC 60,034-1 that the AC motor should be able to endure 1.2 times the rated speed [33]. Therefore, the centrifugal force under 1.2 times the rated speed is set in the FEA software. The stress distribution and the rotor deformation are calculated and depicted in Fig. 3 at 21,600 r/min. Although the maximum stress and deformation are larger than before, the safety requirements are still met. One of the most significant processes in the shaft system design is the rotor dynamic analysis, which has a great impact on the vibration and the acoustic noise of the shaft system [34]. If the motor rotates at the critical speed, the large vibration amplitude may result in collision of the stator and rotor due to the resonant vibration at critical speeds [35]. In order to identify whether the vibration of the rotor is within the acceptance limit, the unbalance response of the shaft system should be analysed by calculating the vibration amplitude at different speeds"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001907_ijfsa.2016010102-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001907_ijfsa.2016010102-Figure4-1.png",
        "caption": "Figure 4. FBD of pendulum",
        "texts": [
            " The IP system comprises of mainly two sub-systems i.e. Cart and Pendulum (Kharola & Gupta, 2013) as shown in Figure 2. A pendulum of mass (m), hinged by an angle \u03b8( ) from vertical axis and mounted on a cart of mass (M). This cart is free to move in horizontal direction with the help of Force (F). The Coefficient of friction acting between cart and ground (b), length of pendulum (L) and Inertia of Pendulum (I). A view of forces acting on Cart and Pendulum are shown with the help of free body diagrams (FBD) as shown in Figure 3 and Figure 4. The FBD are further used for developing governing mathematical equations for IP system. The mathematical equations are derived using Newton\u2019s Second Law (Akole & Tyagi, 2008; Radhamohan et al., 2010) and are shown below separately for each sub-system. Use the following equations for cart evaluation: Use the following equations for pendulum evaluation: I \u03b8 \u03c4=\u2211 (4) \u03b8 \u03b8 \u03b8= +( )1 I NL PLcos sin (5) where, N and P are the interaction forces between cart and pendulum, x is position of cart along horizontal direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure18-1.png",
        "caption": "Fig. 18. Comparison of equivalent strain b",
        "texts": [
            " A detailed comparison between smoothed and unsmoothed simulations is presented in [24], where normals are computed by the consistency approach. In this paper, the influence of the normal calculation is evaluated by comparing normal voting with consistent normals. Fig. 17 presents the contact area. It shows that even though smoothing with consistent normals provides an enhanced description of contact area [24], the use of normal voting gives an even better description of tool geometry and consequently of contact interactions during the entire simulation. Fig. 18 also shows that equivalent strains are more homogenous and have lower values: the maximum of equivalent strain is about 0.64 using consistent normals and about 0.60 using normal voting. This example shows the importance of the accurate description of contact and of enhanced smoothing procedure. The geometries of tool and bar employed in the shape rolling simulation are shown in Fig. 19 \u2013 right. The material is viscoplastic (Norton Hoff law (3) with: K = 158 MPA s0.18, m = 0.18) and sliding contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003707_acsaelm.0c00837-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003707_acsaelm.0c00837-Figure8-1.png",
        "caption": "Figure 8. (a) Schematic and experimental photos for printed robotic arm actuate at a low electric field without prestretch or without attaching passive layers to the elastomer film. (b) Relationship between the opening displacement and the applied electric field for the printed robotic arm.",
        "texts": [
            " It can also do the same for a hollow cylindrical paper weighing 0.2 g without causing damage to the body of the cylinder, as seen in Supporting Information Movie S4. Figure 7b shows the relationship between the electric field and the maximum weight that the device can lift. It is obvious that a piece of elastomer weighing 0.2 g can hold a weight of 1.5 g, which weighs more than 7 times its weight. Let us mention here also that this piece of elastomer is attached with other 3D-printed structures for supporting purposes. Figure 8a is the demonstration of an opening\u2212closing arm fabricated from D20 with a hollow cross-sectional area (1 mm diameter pillars sandwiched between two faces of 0.5 mm layer thickness). The printed arm has a total thickness of 2 mm and can open until a displacement of 30 mm under an electric field of 2 kV/mm, which can allow it to be used as a soft robotic arm with proper strength. This configuration of the arm is 2 https://dx.doi.org/10.1021/acsaelm.0c00837 ACS Appl. Electron. Mater. XXXX, XXX, XXX\u2212XXX I mm thick, which gives a proper strength instead of using thin layers attached to passive layers to actuate at lower electric field and to produce more shape change as used in the literature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000172_j.jfranklin.2019.11.054-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000172_j.jfranklin.2019.11.054-Figure1-1.png",
        "caption": "Fig. 1. System configuration of indoor leader-following.",
        "texts": [
            " Using such a system in an indoor environment in which the GPS is unavailable, a moving reference frame attached on the following robot is always used to define the pose (position and orientation) and trajectory of the leading robot. The advantage of using a moving reference frame is that no global information (such as from GPS) is required and the needed relative information can be measured with low cost sensor sets [3] . In this work, we consider the indoor leader-following of WMR with a system configuration containing a RGBD camera attached on the follower and encoders equipped on each wheel of the leader and follower, as shown in Fig. 1 . The control objective is to drive the follower to follow the trajectory of the leader without explicitly planning any trajectory. Therefore, the development of such a control system for the follower can be decomposed into two sub-problems: trajectory acquisition and controller design. Technically, the acquired trajectory information should be stored in the memory and updated online and then used by the designed controller for real-time motion control of the follower. To obtain the trajectory information, vision is a promising approach in practical applications"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000882_0954405414564405-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000882_0954405414564405-Figure6-1.png",
        "caption": "Figure 6. Analysis of pitch error.",
        "texts": [
            " Considering equation (1), we can see that the input of the C-axis is bounded because the B-axis is BIBO stable. Gear pitch error analysis In this section, the source of pitch error is analyzed, and an instantaneous pitch error formula is deduced. Considering the generating process kinematics, in the case of a hob with one start, each hob tooth penetrates into the next gear gap in the same generating position, removing a chip with the same geometry as in the previous gear gap. As illustrated in Figure 6, P1 is one point on the pitch curve of the hob, and P01 is the matching point on the pitch curve of the workpiece. In the generating process, there should be a series of fixed points that coincide with P1, that is, PC1, PC2, PC3 and so on. The actual point P01 will not be coincident with those points, however, because of the tracking error of each axis and the synchronization error of the EGB. As shown in Figure 6, given that RC is the standard pitch circle radius of the workpiece, m is the module and ZC is the number of teeth of the workpiece, the instantaneous pitch error fpt can be calculated as fpt =RCa= ZCma 2 \u00f02\u00de where a is the tracking error of the C-axis, giving fpt = ZCmEC 2 \u00f03\u00de The gear hobbing CNC hardware platform used in this research is illustrated in Figure 7. It is composed of an ARM9 (EP9315) microcontroller and the DSP TMS320C6713. The complex operation and real-time control are performed by the TMS320C6713 DSP"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002979_tii.2020.2986805-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002979_tii.2020.2986805-Figure3-1.png",
        "caption": "Fig. 3. The configuration and force analysis of vertical insertion. a) Configuration. b) In tension. c) In compression.",
        "texts": [
            " The upper object is connected to the upper manipulator via a spring, using soluble glue for fixation and solvent to dissolve after assembly. The lower object is gripped by the lower manipulator by vacuum absorbing, or pin sealing, or gluing, etc. The task is to insert the upper object into the hole of the lower object with an interference fit between them. The objects are small-sized and thin-walled, which demands highly on force control. The insertion starts with the objects well aligned (coincident axes and no distance between objects) with no contact between them. Fig. 3(a) shows the configuration of the vertical compliant insertion. According to the motion type of insertion or withdrawal, the spring is in tension or compression, with the forces upon the upper object shown in Figs. 3(b) and 3(c). Since in precision assembly, the misalignment is much small (let alone to be compared with the spring\u2019s free length) and we use a straight line to approximate the real curve with very small curvature in modeling the spring configuration. In Fig. 3(b), the spring is in tension and the center of the manipulator deviates from the center of the objects. Label F as the spring force upon the upper object F = k (L\u2212L0) , (1) where L is the current spring length, L0 is the spring length with no load, and k is the spring coefficient. Label Fh as its projection onto the horizontal plane. Supposing the spring force direction is coincident with the spring axis leads to Fh F = kh\u2206xs k (L\u2212L0) = \u2206xs L , (2) where kh is the horizontal stiffness reflecting the rate between the horizontal force and displacement, and \u2206xs is the horizontal center distance between the manipulator and the upper object",
            " Combining the above three equations, the compliance of a tensile spring results in kt = Fh \u2206xt = Mkh kh +M , (8) where kt is the horizontal stiffness with a tensile spring. Considering (5), the horizontal compliance of a tensile spring is 1+ M kh times of insertion without compliant equipment. Since kh relates to the force magnitude, the lowest ratio of horizontal compliance is 1 + M k times. This means that a vertically equipped spring can also enhance the insertion compliance in the horizontal plane. Fig. 3(c) shows the posture where the spring is in compression, and \u2206xs and \u2206xo locate on different sides of the object\u2019s main axis. The upper manipulator\u2019s movement depends on the magnitudes of spring displacement and of object deformation \u2206xc = \u2225\u2206xs \u2212\u2206xo\u2225, (9) where \u2206xc is the motion of the upper manipulator to align the objects and the manipulator. In the case kh equals M, \u2206xc is zero while Fh is non-zero. For this setting, the posture of the spring and the upper object is balanced and the manipulator needs to move a little to break the posture balance and then move back"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002183_2168-9792.1000163-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002183_2168-9792.1000163-Figure6-1.png",
        "caption": "Figure 6: Fabricated Quadcopter model.",
        "texts": [
            " An RF transmitter receives serial data and transmits it wirelessly through RF through its antenna connected at pin 4. The transmission occurs at the rate of 1Kbps - 10Kbps. The transmitted data is received by an RF receiver operating at the same frequency as that of the transmitter. The RF module is used with a pair of encoder or decoder. The encoder is used for encoding parallel data for transmission feed while reception is decoded by a decoder. The assembly and connections of the various electronic components of the Quadcopter is shown in Figure 5. The fabricated quadcopter model is shown in the Figure 6. The total mass of the quadcopter is estimated in Table 2. The total empty mass estimated from the above table is about 901 grams. As the expected payload capacity is considered as 300 grams, the quadcopter should be able to fly with a total mass of around 1200 grams. [10-13] The thrust of the quadcopter [4] is given by the equation T = \u03c0D2\u03c1v\u0394v/4 Where T is thrust in N, D is Propeller diameter in m, \u03c1 is Density of the air \u2013 1.22 kg/m3 Also V = \u0394V/2 Where V is the velocity of air at the propeller, \u0394V is the velocity of the air accelerated by propeller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002796_978-1-4842-5531-5_3-Figure6-12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002796_978-1-4842-5531-5_3-Figure6-12-1.png",
        "caption": "Figure 6-12. Step 5 to obtain the propulsion or allocation matrix",
        "texts": [
            " Very unequal distances, very different angles, or significantly different heights change the propulsion matrix considerably. Nevertheless, if the differences are moderate, this does not influence the operation of the vehicle. See Figure\u00a06-11. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 254 Step 5: Place the coordinate frame of the drone (that is the body frame) and label the axes. This influences the vehicle\u2019s movement and control design. The choice will depend on the test of the remote control levers. See Figure\u00a06-12. Chapter 6 QuadCopter Control with\u00a0Smooth Flight mode 255 Step 6: Relate the rotational frames and their direction of rotation. In this example, you will associate the pitch with the X axis of the drone, the roll with the Y axis, and the yaw with Z. The positive direction of the rotation of the planar axes X and Y is the one in which it moves towards the positive zone of the axes when the vehicle is tilted. The direction of rotation of the Z axis is free, but the right hand rule is usually employed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000175_icems.2019.8922428-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000175_icems.2019.8922428-Figure1-1.png",
        "caption": "Fig. 1. Definition of filling factors for round and flat wires",
        "texts": [
            " A larger proportion of conductor material can lead to an increase in output and higher efficiency by increasing the amount of windings. This means that the same power can be generated in a smaller package of an electrical machine with a higher filling factor. The mechanical filling factor and the electrical filling factor, also known as copper filling factor, are particularly relevant for the design. [3] [4] The mechanical filling factor is calculated as the quotient of wire area to the slot area as can be seen in Fig. 1. The former is the product of the area of the conductors including their coating class and the quantity of individual wires in a slot. The latter corresponds to the slot cross-sectional area after subtracting the insulation of the slot. In contrast to the mechanical filling factor, only the copper content without the insulation is taken into account for the dividend of the copper filling factor. For the divisor the total slot cross-sectional area including the slot insulation is considered. [3] Since the program is supposed to calculate filling factors for flat wires as well, the formulas explained above must be adapted accordingly. In this case the wire area is no longer calculated based on the diameter of the conductor, but from the product of the corresponding length and width of the flat wire and the subtraction of the edge radii. The formulas for calculating the two filling factors for round or flat wires are presented in Fig. 1 below. III. PACKING DENSITY The filling factor is predominantly determined by the quality of the layer structure, the so-called packing density, within a winding with round conductors. The larger the packing density, the smaller is the space between three conductors. The size of the packing density is affected on the one hand by the pattern of the layers of a winding and on the other hand by the conductor geometry. [2] In the pattern of the layer structure of the winding, a distinction is made between the random, the accurate positioning and the orthocyclic winding, shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000803_j.actaastro.2014.04.009-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000803_j.actaastro.2014.04.009-Figure1-1.png",
        "caption": "Fig. 1. Model of a planar dual-link space robot (\u00a9 JSASS). \u22123 \u22122 \u22121 0 1 2 3",
        "texts": [
            " In Section 4, a combination of the MBSC and the input-shaping techniques, which is collectively referred to as IS-MBSC, is explained. In Section 5, the dependency of the vibrational frequency on the link angles is investigated experimentally and is then formulated in the form of a fourth-order function of the link angles using the response surface method. Experimental results are also presented in order to demonstrate the effectiveness of the proposed method. Finally, conclusions are presented in Section 6. Fig. 1 shows a schematic diagram of a planar space robot consisting of a dual-link manipulator connected by revolution joints without a flexible appendage. Assuming that the orientation of the main body is \u03b8, the angle of the first link (angle 1) is denoted by \u03d51, and the angle of the second link (angle 2) is denoted by \u03d52. The masses of the main body, the first arm, and the second arm are denoted by m0, m1, and m2, respectively, and correspondingly, J0, J1, and J2 are the moments of inertia of the \u03c6 1 [rad] main body, the first arm, and the second arm, respectively, around their respective centers of mass. The notation for the dimensions and joint angles of the model is given in Fig. 1. Given that the initial total angular momentum is zero and that no external torque or force affects the system, the law of angular momentum conservation holds. Accordingly, the angular velocity of the main body of the space robot is represented by functions of the two link angles h1 and h2, the angular velocities of which are as follows: _\u03b8 \u00bc h1\u00f0\u03d51;\u03d52\u00de _\u03d51\u00feh2\u00f0\u03d51;\u03d52\u00de _\u03d52 \u00f01\u00de We assume that the state vector x and the control input u are given as x\u00bc \u00bd\u03d51;\u03d52; \u03b8 T and u\u00bc \u00bdu1;u2 T \u00bc \u00bd _\u03d51; _\u03d52 T , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure4-1.png",
        "caption": "Fig. 4. Conventional 6S-10P C-Core SFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " Flux distribution of 12S-10P E-Core SFPMM is shown in Fig. 2(b). Analysis reveals that despite of reduce in PM usage, still there are significant flux leakage. A similar performance evaluation of conventional 6S-10P SFPMM with E-Core and C-Core is carried out in [12] that reduce PM usage and enhance slot area. Moreover, its performance is compared with conventional 12S-10P E-Core SFPMM. Fig. 3(a) and Fig. 3(b) shows 2-D cross sectional view and flux distribution of 6S-10P E-Core SFPMM whereas Fig. 4(a) and Fig. 4(b) shows 2-D cross sectional view and flux distribution of 6S-10P C-Core SFPMM. Analysis reveals that conventional 6S-10P E-Core SFPMM and 6S-10P C-Core SFPMM retain same PM volume as that of 12S-10P E-Core SFPMM and increased slot area, however author fails to compensate effects of leakage flux. Until now, many researchers tried to reduced PM as much as possible and suppress flux leakage but unfortunately both effects are not considered at a time. In this paper, alternate Consequent Pole SFPMM (CPSFPMM) with partitioned PMs are introduced which further reduce PM usage and suppress flux leakage completely"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002213_075045-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002213_075045-Figure4-1.png",
        "caption": "Figure 4. The structure of cylinder head to install FBG sensors.",
        "texts": [
            " The load hydraulic cylinder 14 is to apply load on the test hydraulic cylinder. Overflow valves 18 and 22 adjust back-pressures of the inlet and outlet chambers, so as to regulate the load size. In the experiment, load size will be kept the same in the instroke and outstroke process. Pump 24 is only used to maintain oil circulation, and the pump output pressure should be adjusted to the minimum through overflow valve 23. The structure design scheme of the cylinder head to install FBG sensors is shown in figure 4. FBG sensors are embedded in the bottom of the seal groove to measure the contact stress of the seal. It should satisfy two requirements: (1) to not affect the normal operation of reciprocating hydraulic cylinders; (2) light intensity loss in the process of optical transmission should be as little as possible. As shown in figure 4, the cylinder head is designed as three parts: gland, backing ring and cylinder head., where the backing ring is used to monitor the contact strain and temperature of the seal. The cylinder head with FBG sensors is shown in figure 5. In order to compare the contact strain changes of the seal in failure condition, the contact surface of a normal seal was worn to simulate particle scratch, which is shown in figure 6. As shown in figure 7, pressure and displacement signals are collected by NI9021 acquisition card, pressure sensor is AK4a and displacement sensor is SMW-LX-08-800- V2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002253_1464419316660930-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002253_1464419316660930-Figure9-1.png",
        "caption": "Figure 9. Scheme of the damage introduced to the rings: (a) damage position on the outer ring, (b) detailed view of the damage introduced to both rings.",
        "texts": [
            " Firstly, results regarding the simulation of a single-row angular-contact ball bearing are shown, specifically the REXNORD ER16K bearing is simulated; then, the response of a NJ 305 cylindrical roller bearing is shown. The dimensions of the two REBs are shown in Table 5. In both cases, three different scenarios regarding the health condition of the bearing are considered: a bearing without damage, a case with local damage in the outer ring and another one with local damage in the inner ring. In the last two cases, the damage is introduced as changes on the surfaces of the damaged bodies. Figure 9 shows that the damaged surfaces have been defined by means of two smooth curves. These cubic curves are defined by the angular length of the damage, , and the damage depth, h. Moreover, the properties of the steel have been defined as the same for both case studies: modulus of elasticity E of 207 GPa, Poisson\u2019s ratio of 0.3 and density of 7830 kgm 3. The model is verified and validated according to the guidelines provided by ASME63 by the use of theoretical data, experimental data and data obtained from simulations of another model in the literature, as explained throughout this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure15-1.png",
        "caption": "Fig. 15 Flank topographic deviations between the theoretic and produced tooth surfaces: (a) solid model (STL file) produced by VERICUT and (b) flank topographic deviations",
        "texts": [
            " Because only a cutter is employed, the right and left flanks are cut in the upper and lower positions (verified by SOLIDWORKS and illustrated in Fig. 13), but in contrast to the traditional machine, the CNC machine allows freeform flank modification. These coordinates can be used to program the NC data needed by the VERICUT software to confirm the correctness of machine positioning in the manufacturing stage. Figure 14 shows the cutting simulation for the pinion using VERICUT. The resulting tooth surfaces are saved as an STL (stereolithography) file and later compared with the theoretic tooth surfaces. Figure 15 shows the flank topographic deviations between the theoretic and produced tooth surfaces. Because of the resolution of the triangle meshes (an interpolation tolerance of 0.05 mm), the sum of the squared errors is 38; 866 lm2 and the tooth thickness error is \u00fe22:6 lm This result, which contains discontinuous areas on the surfaces generated, reflects the error caused by VERICUT simulation. Nevertheless, for the most part, the results confirm that the tooth surfaces produced are correct. The high productivity of the DICC method for SBGs comes from the usage of large circular plates with many cutter teeth and a cutter tilt that enables lengthwise crowning to produce convex teeth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003583_biorob49111.2020.9224389-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003583_biorob49111.2020.9224389-Figure3-1.png",
        "caption": "Fig. 3. A robotic patient positioner with a patient chair and the soft immobilization device attached.",
        "texts": [
            " In our study, the novel soft immobilization is designed based on a fluid-driven origami-inspired artificial muscles (FOAM) concept [9]. This is an easy to make and very cost-effective design. To focus the effort on developing and evaluating the soft immobilization device, we used a standard 6 DOF parallel robotic base with a simple controller in this study. In addition, a chair top was integrated with this soft immobilization device to achieve a clinical required position control system. The system developed in this study (Fig. 3) is comprised of a parallel robot base, a patient chair with a soft immobilization device attached to it, and an optical motion tracking system. This system was developed to allow immobilization of a patient, while providing accuracy positioning of the patient. The following requirements were considered when developing a robotic patient positioner combined with a soft immobilization device for a compact proton therapy system. 1) The immobilization device needs to be a) thin, i.e., < 5 cm to allow for positioning close to the radiation beam nozzle",
            " The immobilization unit was connected to a vacuum pump with a regulator and air pressure sensor attached. The pressure was regulated to vary the pressure difference between the device pressure and atmospheric pressure, while the load cell was monitored. For each of the 80 different pressure differences tested, the force registered by the load cell as shown in the pressureforce plot in Fig.6. A maximum force of (150\u00b14)N was achieved at (99\u00b11) kPa. A linear relation between pressure and force was observed with a small hysteresis effect. The robotic patient positioner developed in this study (Fig.3) consists of a Gough\u2013Stewart [23], [24] platform and a patient chair. A commercial Stewart platform (DOF Reality, Corp.) designed for driving and flying simulations was adapted as the base of the patient positioner. The six rotary actuators of the system work in pairs connected to a motor driver each controlled by a microcontroller (Arduino UNO, Italy). For this study the patient chair mounted on top of the Stewart platform was designed specifically for breast cancer treatment. The chair has height adjustable arm supports for lateral arm placement that provides support under the arm and in the armpit region to prevent slouching"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001207_1077546314562621-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001207_1077546314562621-Figure1-1.png",
        "caption": "Figure 1. Rolling element ball bearing geometry.",
        "texts": [
            " In the case of data acquisition, the window of time series of data has to be minimally at least one revolution of the shaft to cover all possible defect frequencies. Table 1 lists the actual (1)\u2013(4) and approximate (5)\u2013(8) relations for the bearing defect at FRESNO PACIFIC UNIV on January 9, 2015jvc.sagepub.comDownloaded from frequencies: where frm, dpitch, dball, n and represent frequency of rotation, pitch diameter, ball diameter, number of balls and the contact angle respectively as highlighted in Figure 1. Among the bearing defect frequencies, the cage defect frequency is the lowest in the frequency band. Given the speed of rotation of the shaft as frm, sampling rate sr, the number of samples required to cover one complete revolution of the shaft for the cage defect frequency are given by (9): NCD \u00bc 1 fCD sr: \u00f09\u00de Automated bearing failure detection considers the minimum window size against the cage defect frequency to incorporate the impact of machine dynamics over one complete revolution. It is because the vibration pattern against the fault varies with the location of the fault points during the course of rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure1-1.png",
        "caption": "Fig. 1 The schematic of machining conical thread surfaces with lathe center displacement",
        "texts": [
            "eywords Conical worm . Lathe driver . Ellipse . Pitch The worm shaft is driven with the help of the driving pin through the lathe fork because of lathe center displacement (Fig. 1). As a result of the shifting of the worm shaft by half cone angle, the path curve of the driving pin will be an ellipse path instead of a circle on the perpendicular plane to axis [1]. The peripheral speed of the spindle is constant, but due to the ellipse path, the radius is constantly changing as a function of time. That is why the angular velocity and the angular rotation are also changing, and these cause pitch fluctuation during the manufacturing process of conical worms. Our objective was to determine the geometrical shaping of the driving pin in order not to have pitch fluctuation [2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure3-1.png",
        "caption": "Figure 3. Finite-volume partitioning (zoomed).",
        "texts": [],
        "surrounding_texts": [
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nThe drum and beater speeds have also been parameterized. The physical parameters of the air were set:\n Density \u2013 1.225 ,\n Viscosity \u2013 1.79 kg/(m\u2219s),\n Acceleration of gravity 9.81 .\nNumerical-analytical algorithm was implemented using two transport equations for turbulent\ncharacteristics (k-\u03b5 model) [18].\nTo obtain the fields of air velocities in the housing of the final threshing device, this problem was\nsolved numerically by the method of finite volumes [17].\nTo implement the numerical-analytical algorithm, the following conditions were set, boundary\nconditions:\n1. Fan speed in rpm is set. 2. It section 1 (Figure 2), air is drawn in, the pressure is 0 Pa. Here and below, the\npressure is specified as the difference between atmospheric pressure.\n3. In section 3, Figure 2, the air flow is released, the pressure is 0 Pa. 4. In section 2 (Figure 2), the air flow rate is set to 20 m/s. 5. Sticking conditions are set on the other walls of the casing, the casing and the walls of\nthe fan.\nAll geometrical dimensions, drum rotation speed and air flow drawing speed were parameterized to\nmake it possible to study the dependence of the sought parameters on these characteristics.",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nFigures 2, 3 show a finite-volume partition of the internal air domain of a field machine.\nTo obtain the distribution fields of flow rates in the body of the field machine, a series of calculations was carried out at different geometric parameters and drum rotation speeds.\nFigures 4 - 8 show the contours of the velocity fields and the velocity vectors for the initial geometric parameters (Figure 1). The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 36.2 m/s. On the left side of the drum, an",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nincrease in the flow rate is observed, due to the narrowing of the space between the drum and the housing wall.\nA slight discrepancy in the display of velocities in the graphs of displaying fields and vectors is due\nto the peculiarities of displaying and visualizing numerical results."
        ]
    },
    {
        "image_filename": "designv11_30_0003978_978-3-319-14705-5_12-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003978_978-3-319-14705-5_12-Figure1-1.png",
        "caption": "Fig. 1 Structure of spherical robot BHQ-1 (1: motor 1, 2: motor 2, 3: mass, 4: shell, 5: camera, 6: bearing, 7: controller & battery, 8: hollow axle)",
        "texts": [
            " established the dynamic model of a spherical robot Omnibola with Newton-Euler equations and compared its actual motions with the simulated ones through experiments [18]. In the following of this chapter, a brief introduction of spherical robot BHQ-1 will be given first, and then one motion planning method based on kinematics and one motion planning method based on dynamics will be introduced separately. BHQ-1 is the first kind of spherical robot designed by our lab and the first prototype was implemented in 2001 [8], which is designed for the exploration of unmanned environments. As shown in Fig. 1, BHQ-1 is mainly composed of two motors, one hollow axle, one mass, one camera, one controller and battery combination, and one ball-shaped shell. In Fig. 1 frame {XbYbZb} is a body frame attached to the hollow axle and its origin is coincident with the geometric center of the sphere. The hollow axle connects with the shell through two ball bearings at the two ends and serves as a chassis or frame to install other components, so the outer shell can rotate around the axis of the hollow axle freely and the camera installed on the hollow axle can keep a relatively steady posture no matter BHQ-1 is moving or static. Motor 1 is installed on the hollow axle but its output axle is fixed to the shell, so its rotation can result in the displacement of the mass along Yb direction",
            " Installed on the hollow axle the camera is used to take pictures of environments which can be transmitted to a remote control center through a wireless image transmission system. According to the received pictures an operator can not only observe the environment but also control the motion of the spherical robot through a joy stick. The motion principle of the spherical robot is that the rotations of motor 1 and motor 2 make the mass rotate about axes Xb and Yb respectively and result in the displacement of the center of gravity of the whole system, which produces a displacement moment to counteract the friction moment and makes the robot move. As shown in Fig. 1, when motor 1 rotates and motor 2 keeps still, the mass, the hollow axle, the controller and battery combination, and motor 2 will rotate about the axis of the hollow axle. If the angle displacement \u03b8 \u2265 \u03b80 (\u03b80 is the angle displacement of the mass to balance the moment caused by static friction), the robot will move forward or backward. Because the moment caused by dynamic friction is less than that caused by static friction, the mass will stay at a position where the angle displacement of the mass is less than \u03b80"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003329_aim43001.2020.9158972-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003329_aim43001.2020.9158972-Figure18-1.png",
        "caption": "Figure 18. Relationship between time and crawling distance",
        "texts": [
            " Paying attention to the graph to which this correction is applied, the speed becomes negative after about 140 s. This was not owing to an error, but it was confirmed from a video image of the crawling state of the robot that the PEW-ROV was traveling backward because it was caught when passing through a 90\u00b0 bending pipe. In addition, it was confirmed that the PEW-RO V slipped and moved backward from about 240 s to 280 s. Therefore, it is considered that such a sudden backward movement can be accurately represented. Fig. 18 shows graphs of the estimated crawling distance when this correction is applied and when it is not applied. Table 4 shows a comparison between the actual crawling distance and the estimated crawling distance, and a comparison with the previous study [13]. When the correction was not applied, an error of approximately 428 % occurred, but when the correction was applied, the error was only about 1.42 %. In the previous study, the error for the moving distance was 7.60 %, and in this experiment, the error for the moving distance was 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001853_hnicem.2015.7393247-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001853_hnicem.2015.7393247-Figure2-1.png",
        "caption": "Fig. 2 Moments of inertia of the quadrotor [9]",
        "texts": [
            " The variables of the quadrotor are: x = inertial position of the quadrotor alongside xi in Ri y = inertial position of the quadrotor alongside yi in Ri z = inertial position of the quadrotor alongside -zi in Ri u = frame velocity measured in xv in Rv v = frame velocity measured in yv in Rv w= frame velocity measured in zv in Rv = roll angle with respect to Re \u03b8 = pitch angle with respect to Re \u03c8 = yaw angle with respect to Re p = roll rate measured alongside xv in Rv q = pitch rate measured alongside yv in Rv r = yaw rate measured alongside zv in Rv Shows that: m = total mass of the quadrotor g = gravitational constant Jx, Jy, Jz = moments of inertia around x, y, and z computed by assuming the center of the quadrotor is a sphere with mass M and radius R, and the mass m of the motors l distance from the center of the quadrotor shown in Figure 2: The moment of inertia for a sphere is[9] 2 5 5 It follows that: 2 5 2 6 2 5 2 7 2 5 4 8 B. Forces and Torques With the kinematic and dynamic equations set. The forces and torques acting in the quadrotor will be defined. The relationship of lift generated by the motors and the gravitational pull is established. [9] Upward Force: Quadrotor thrust is the sum of the thrust by each propeller. 9 Rolling Torque: Torque produced by increasing the right motor\u2019s thrust and decreasing the left motor\u2019s thrust and viceversa while keeping the front and back motor\u2019s thrust the same"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure15.8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure15.8-1.png",
        "caption": "FIGURE 15.8 Perspective view of concentric square transmission line.",
        "texts": [
            " Since there are only two variables involved, these structures may be zoned in the same manner as 2d rectangular problems, but the interpretation of the drawing is very different. For example, consider the example used in Section 15.2 (Figures 15.2 and 15.6). These figures described the cross section of an infinitely long (in z) pair of concentric squares. Figure 15.7 shows the various ways this structure may be zoned, either using the full cross section (a) or taking advantage of various axes of symmetry (b), (c), and (d). Figure 15.8 shows a perspective view of the same concentric square transmission line. An important point to keep in mind here is that Figure 15.8 does not show the entire structure described. It just shows a section (in z) of the structure. The structure is infinitely long in Z, and we have been dealing only with a cross section of it. Since such a structure with no Z dependence (i.e., uniform) can be described entirely by voltages and fields that are only functions of X and Y, a 2d analysis finds everything there is to know about the structure. 15.3 Axisymmetric Structures 337 Voltages, electric field components and energy density are functions of X and/or Y only and are valid for any value of Z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001131_j.amc.2014.01.138-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001131_j.amc.2014.01.138-Figure3-1.png",
        "caption": "Fig. 3. The ortho\u2013parallel manipulator in configurations corresponding to ideals W9;W10 and W85 illustrated with computer program Alaska .",
        "texts": [],
        "surrounding_texts": [
            "We represent the orientation and the corresponding rotation matrix using Euler parameters. Let a \u00bc \u00f0a0; a1; a2; a3\u00de and define the matrix Rby R\u00f0a\u00de \u00bc 2 a2 0 \u00fe a2 1 1 2 a1a2 a0a3 a1a3 \u00fe a0a2 a1a2 \u00fe a0a3 a2 0 \u00fe a2 2 1 2 a2a3 a0a1 a1a3 a0a2 a2a3 \u00fe a0a1 a2 0 \u00fe a2 3 1 2 0B@ 1CA: If a belongs to the unit sphere a 2 S3 R4; jaj \u00bc 1 then R\u00f0a\u00de is orthogonal with det\u00f0R\u00f0a\u00de\u00de \u00bc 1 and we get a polynomial representations for rotation matrices. Note that R\u00f0a\u00de \u00bc R\u00f0 a\u00de and from this it follows that sometimes we may have 2 or more distinct subvarieties of the configuration space which nevertheless correspond to the same physical positions of the mechanism. Perhaps a simple example clarifies the situation. Consider a single rigid body with a constraint 2a2 3 1 \u00bc 0. In this situation I \u00bc hjaj2 1;2a2 3 1i \u00bc I1 \\ I2 \u00bc h2a2 0 \u00fe 2a2 1 \u00fe 2a2 2 1; a3 1= ffiffiffi 2 p i \\ h2a2 0 \u00fe 2a2 1 \u00fe 2a2 2 1; a3 \u00fe 1= ffiffiffi 2 p i: Let us further denote V \u00bc V\u00f0I\u00de and Vj \u00bc V\u00f0Ij\u00de. Now interpreting R as a map R : S3 ! SO\u00f03\u00de then it is clear that R\u00f0V\u00de \u00bc R\u00f0V1\u00de \u00bc R\u00f0V2\u00de; although Vj V and moreover V1 \\ V2 \u00bc ;. From the point of view of analysis of the configuration space it is evident that we can simply take one of the Vj and ignore the other. Since this situation occurs frequently it is convenient to have a notation for it. Suppose we have m rigid bodies whose orientations are constrained to be in some variety V \u00f0S3\u00dem. Then the map R induces in a natural way a map R\u0302 : V ! \u00f0SO\u00f03\u00de\u00dem. Now suppose that Vj V are some subvarieties. Definition 3.1. V1 and V2 are physically same, if R\u0302\u00f0V1\u00de \u00bc R\u0302\u00f0V2\u00de. In this case we write V1 V2. Moreover if V1 is a proper subvariety of V2 then we write V1-V2. Hence the goal in the analysis is to find the \u2018\u2018smallest\u2019\u2019 or \u2018\u2018most convenient\u2019\u2019 W such that W-V ."
        ]
    },
    {
        "image_filename": "designv11_30_0001534_0954410012472292-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001534_0954410012472292-Figure1-1.png",
        "caption": "Figure 1. A flying-wing type research UAV.19 UAV: unmanned aerial vehicle.",
        "texts": [
            "comDownloaded from The later section deals with the trim and stability analysis of the damaged flying-wing UAV. Then, the dynamic analysis of the damaged UAV by conducting numerical simulations with its longitudinal and lateral/directional models is discussed. Lastly, conclusions and future works are given. A flying-wing type single jet UAV was developed for modeling and analysis of the damaged airplane. The partial right-wing damage is assumed to have a loss of area moments of 22%, 60%, and 78.4%. Figure 1 shows the geometry of the flying-wing type UAV considered in this research, whose body/span lengths and sweep angle are 1.883m/1.98m and 43 , respectively. The partial wing loss causes the UAV to reduce its total mass. Figure 2 represents the trend of the mass decrease with respect to the percentage of wing damage, where the total mass reduces to 12% when the wing area moment is damaged by 78.4%. This is found from using CATIA measure inertia function, Table 1. Test condition of wind tunnel test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000608_jae-141776-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000608_jae-141776-Figure8-1.png",
        "caption": "Fig. 8. Variation of axial component of a magnetic field with angular displacements of the rotor. (a) PM bearing with axial polarized rings. (b) PM bearing with radial polarized rings. (c) PM bearing with perpendicularly polarized rings.",
        "texts": [
            "02 0 5 10 15 1 2 3 4 5 x 105 Axial Displacement [m]Angular Displacement [deg] R ad ia l C om po ne nt , H r [ A /m ] 1.5 2 2.5 3 3.5 4 4.5 x 105 (a) (b) For circular rings, due to symmetry, HX = HY = Hr, radial component of the magnetic field. A variation of the axial component of the magnetic field with radial displacements (0.5, 0.75 and 1 mm) of the rotor for its different axial positions in PM bearing with axial, radial and perpendicularly polarized rings is presented in Fig. 7 (axial magnetic field component is calculated at the centre of the left face of the inner ring). Figure 8 represents the variation of the axial component of the magnetic field with angular displacements (5, 7.5, 10, 12.5 and 15 deg.) of the rotor for its different axial positions in PM bearing with axial, radial and perpendicularly polarized rings (axial magnetic field component is calculated at the centre of the left face of the inner ring). Results shown in Figs 7 and 8 demonstrate that the effect of radial displacement of the rotor on the axial component of a magnetic field is least, whereas the effect of angular displacement is significant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001059_robio.2013.6739697-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001059_robio.2013.6739697-Figure4-1.png",
        "caption": "Fig. 4. Grasp postures for ellipsoid and elliptic cylinder",
        "texts": [
            " The following posture is excepted evaluation, and examples are shown in Fig.3. \u2022 There is a collision between a finger and the surface which is not target quadric surfaces \u2022 Target objects are sharper than a threshold Quadric surfaces whose parameters u, v and w are positive are selected from quadric surfaces. For selected ellipsoids, three principal axes r1, r2 and r3 are calculated. The one axis is chosen as the hand approaching axis from these and the one axis is chosen as the gripper closing axis from remaining two axes, as shown in Fig.4(a). The approaching axis has a variation of 6 directions which are positive and negative directions of principal axes. The gripper closing axis has variation of 4 directions as rotation by 90 degrees. Then, a ellipsoid has 24 candidates of grasp postures. The system selects that parameters u, v and w satisfy the conditions as cylinder from quadric surfaces. For selected the elliptic cylinder, two principal axes r1 and r2 of ellipses of the end faces are calculated. The one axis is chosen as the approaching axis from these and the other one is as the gripper closing axis. The approaching axis has variation of 4 directions which are positive and negative directions of the two axes. The gripper closing axis has variation of 2 directions as rotation by 180 degrees. As shown Fig.4(b), upper point, middle point and lower point of the cylinder are set as grasping points. Let l be a length between the end faces of the cylinder. The upper point and the lower point are defined as the position that is moved to +l/4 and \u2212l/4 from middle point respectively. Then, the elliptic cylinder has 24 candidates of grasp postures. The evaluation method of the grasp stability is categorized in two: point contact and surface contact. The point contact is realized using a hand whose finger-tip is made of a rigid material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.76-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.76-1.png",
        "caption": "Figure 4.76 Rotary motion; belt and pulley coupled",
        "texts": [
            " TACC at load = (2736)(62.8) = 171 820 g cm; a 2 : 1 safety margin for the gearhead System Design 237 The belt and pulley drive is similar to the gearhead drive in that there is a ratio between the motor and the load. In this case, N is the ratio of the load pulley diameter (DLP) to the motor pulley diameter (DMP). N = DLP/DMP In addition, the weight of the belt must be reflected to the motor as an inertia, especially if the belt is long or of a high mass material such as a chain link belt. See Figure 4.76. System Data: JM = motor inertia (g cm s2) JL = load inertia (g cm s2) JMP = motor pulley inertia (g cm s2) JLP = load pulley inertia (g cm s2) JS = shaft inertia (g cm s2) TL = load torque (g cm) TF = friction torque (g cm) (seen at motor pulley) BW = belt weight (g cm) Belt weight reflected to the motor as an equivalent inertia: JB = ( BW 980.6 ) ( DMP 2 )2 Motion: Position: \u03b8M = N\u03b8L (rad) Velocity: \u03b8 \u2032 M = N\u03b8 \u2032 L (rad s \u22121) Acc/Dec: \u03b8 \u2032\u2032 M = N\u03b8 \u2032\u2032 L (rad s \u22122) 238 Electromechanical Motion Systems: Design and Simulation At Motor: JT (total inertia) = JM + JL/N 2 + JMP + JLP/N 2 + JB (theoretical) (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002802_lra.2020.2970953-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002802_lra.2020.2970953-Figure3-1.png",
        "caption": "Fig. 3. Related mechanical structures and settings on practice lathe.",
        "texts": [
            " w\u0302s = 1 N N\u2211 n=1 w\u0302n s . (12) III. ILLUSTRATIVE EXAMPLES To validate the accuracy and effectiveness of the proposed DBI scheme, a series of experiments are performed. Details of the experiment setup and collected dataset are given as follows. Results of DBI models\u2019 comparison are also discussed below. The alloy-6061 vehicle wheel with the radius 17 inches and mass 12 kilograms under G40 in the balancing quality grade from ISO 1940/1 is used to conduct the final machining on a practical lathe as Fig. 3, including related structures of the wheel, spindle and hardware of the DBI scheme. As displayed in the red solid square of Fig. 3, the spindle box, the most critical component, rotates a wheel fixed by the clamper through the belt-driven spindle. These heavy components with the high stiffness hide the moment of inertia of rotating wheels. Two accelerometers and one laser device are respectively installed on the surfaces of spindle box and at the place close to the rotating spindle, to collect required data when the lathe wheel finishes the wheel machining, as depicted in the left-hand side of Fig. 3. The sampling rate, the data collection period T, and the rotation speed for Fs are respectively set as 10000 Hz, 30 seconds, and 1920 RPM; thus the data length J of Fs is 300000. Authorized licensed use limited to: University of Exeter. Downloaded on May 05,2020 at 20:32:49 UTC from IEEE Xplore. Restrictions apply. For attaining the most comprehensive learning experience, AD and WE need to collect diverse types and large volumes of training data. To diversify the classes of the labeled samples, one unbalanced wheel with {\u03b8 = 0\u00b0,w = 25g} is repeatedly collected for expanding various symptoms of unbalanced wheels via attaching balancing weights on different positions, as shown in the right portion of Fig. 3. The unbalanced positions are averagely divided into 12 intervals, which means the labels of \u03b8s range from 0\u00b0, 30\u00b0, 60\u00b0, \u2026, to 330\u00b0 and each condition is executed twice to obtain a total of 24 samples for \u03b8s; while the unbalanced weights of 25 labels vary from 0g, 1g, 2g, 3g, \u2026, to 24g via attaching different balancing weights on \u03b8= 0\u00b0 or 180\u00b0, and each condition is executed three times to obtain the total 76 samples (including the original wheel itself with w = 25g) in total for ws.To increase the sample number, the segmentation number N is defined as 40, so that each Fs is segmented into 40 parts of fns and each length of fns would be 7500"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003100_j.matpr.2020.04.476-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003100_j.matpr.2020.04.476-Figure4-1.png",
        "caption": "Figure 4. FEA Analysis for Uniform Porous Structure",
        "texts": [
            "7 mm x 12.7 mm x 12.7 mm as per ASTM A370 Standard. 4. FEA Analysis A.S. Babu et al. / Materials Today: Proceedings 24 (2020) 1561\u20131569 1565 Table 2. Numerical Results of FEA analysis S. No. Porosity (%) Von- Mises Stress (MPa) Deformation (mm) Mesh 1. 10 103.25 0.02225 238298 2. 20 112.58 0.10079 248224 3. 30 142.14 0.10709 248893 4. 40 194.99 0.12469 228453 5. 50 205.25 0.13125 228453 6. 60 208.08 0.09437 224958 7. 70 214.74 0.10561 228467 8. 80 221.94 0.10716 224862 9. 90 230.18 0.13929 235128 Figure 4 and 5 shows FEA analysis for respective uniform and gradient structures. From table 2 where numerical results are tabulated, it is clear that Von-mises stress is inversely proportional to compressive strength. Structure with 90 % porosity is having maximum porosity as well as von-mises stress. Structure with 10 % porosity is having least von-mises stress. So to have a mixture of good strength as well as porosity of the structure, model with 60 % porosity is selected for fabrication. The CAD model of the porous structures generated using K3DSurf v0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002190_tmag.2016.2589924-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002190_tmag.2016.2589924-Figure10-1.png",
        "caption": "Fig. 10. ODS trials (accelerometer positions).",
        "texts": [
            " In order to validate the simulation results obtained by the radial pressure of the electromagnetic model, an ODS was carried out with the Pulse LabShop software. The aim was to obtain the spatial deformation of the stator yoke for a given frequency. This was done by running several vibration measurements along the stator circumference: ten points considered. Radial vibrations are recorded with the aid of four accelerometers, besides a reference accelerometer. These three accelerometers (1-3 in Fig. 10) are sequentially displaced on the mesh defined for the ODS. data acquisition, and postprocessing done in Pulse Labshop software. The machine was run at a constant rotation speed (Ns = fs /p) under load I = 4 A (25% rated current). The goal was to validate the pressure line of order 2 at 2 fs observed in the FE results. Fig. 11(a)\u2013(d) presents the deflection shape observed at different times. A mechanical deformation of order 2 can be seen (ovalization) at 2 fs . It appears that this vibration wave of order 2 is a progressive rotating wave whose origins have been described by ACHFO and the convolution approach links to the rotating wave of the flux density (H 5 and H 7)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002253_1464419316660930-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002253_1464419316660930-Figure2-1.png",
        "caption": "Figure 2. 3D view of the shape of the cage: (a) cage for a ball bearing; (b) cage for a roller bearing.",
        "texts": [
            " For roller bearings, another approach is followed when defining the value of 58 \u00bc A\u00fe B u\u00f0 \u00de exp C u\u00f0 \u00de \u00feD \u00f010\u00de where A, B, C and D are coefficients for a particular lubricant.58 The sum of the normal and tangential forces, leading to the total RE\u2013ring contact force, creates a moment, which is calculated as the cross product between the vector from the GC of each part to the contact points and the contact force. Rolling element\u2013cage contact. In this work, the cage is considered as a body with Z holes for avoiding contact between the REs. These holes have a cylindrical form for ball bearings and take the form of prisms for roller bearings, as illustrated in Figure 2. The contact loads between the REs and the cage holes are calculated using Hertz theory as explained above. Ring\u2013cage contact. The contact between the rings and the cage is simplified by considering that a nonlinear spring connects the centre of the cage and the centre of the rings. This nonlinear spring is formed as a linear spring with a gap in which there is no contact force. Figure 3 shows the relation between the radial distance between the centres rr and the radial components of the contact force as well as the relation between the axial distance between the centres ra and the axial component of the contact force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure8-1.png",
        "caption": "Figure 8. Velocity vector contours for standard dimensions in the Y direction.",
        "texts": [],
        "surrounding_texts": [
            "An optimization calculation was carried out to assess the effect of the hull geometry. The original body geometry (see Figure 1) has been changed by moving the top down 70 mm - Figure 9. Figures 10 - 12 show the contours of the velocity fields and the velocity vectors for the modified housing shown in Figure 9. The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 31.2 m/s. For this geometry, a more even distribution of the air velocity fields is observed. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure10-1.png",
        "caption": "Fig. 10. A 3D car-like type wheeled mobile manipulator with a spherical two-link manipulator.",
        "texts": [
            " Our steering law relied on the geometric human intuition, whereas the one in [4] used the gradient of the configurationdependent manipulability that is not easily predictable from task to task. Fig. 9 shows the evolution of the error and state of the WMM as time goes. The error level remains small but not as good as that from our scheme. As mentioned before, large motions of the mobile platform and the arm joints are observed at the beginning in order to follow the trajectory without switching the moving direction. The steering angle is sometimes saturated like ours. Next, we conduct a 3D trajectory tracking task using another car-like WMM having a spherical two-link arm, shown in Fig. 10(a), whose equivalent virtual manipulator is shown in Fig. 10(b). We apply the proposed inverse kinematic algorithm addressed in section 4. There is a 1-DOF redundancy because the WMM has four active virtual joints in the 3D operation space. As before, the manipulability measure is used for the redundancy resolution, and the heuristic steering law in (21) is used. The numerical values of the wheel-base and link lengths are assumed to be the same as those of the previous planar WMM. The control gains are also set to be the same as previous. The initial conditions are chosen as 1 2 3 [ (0) (0) (0) (0) (0) (0)] and (0) 0[0 0 0 0 / 2 / 2] T p p T x y \u03c6 \u03b8 \u03b8 \u03b8 \u03b3\u03c0 \u03c0= =\u2212 for which the end-effector position is initially located at ( , , ) (2,0,1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure2-1.png",
        "caption": "Fig. 2. A section view of the recirculating hollow-shaft structure.",
        "texts": [
            " Finally, an optimized design of the recirculating hollow-shaft cooling structure will be performed in Section III and the advantages of the optimized structure will be highlighted by comparing the CHTC, which is estimated using CFD. 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 3511 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. II. ANALYSIS OF HOLLOW-SHAFT A rotor cooled by a recirculating hollow-shaft structure is shown in Fig. 2. The rotor is dragged by the prime mover. Heat is generated by passing current to heating pipe, which is inserted on the outer edge of the rotor. In order to measure the average temperature of the wall, temperature sensors are installed on the outer edge of the rotor and near the hollowshaft. Since the heating pipe and temperature sensor are installed on the rotor and rotate with it, the wires on them are connected to the outside through a slip ring. Due to the rotor has slip ring parts at the front and is limited by the size of the bearing, the stationary cooling tube used to introduce the oil is longer and the tube diameter is thinner",
            " The maximum CHTC at the corner of the blade, the local CHTC of the blade area can reach 2038W/m2/K, and the average CHTC of the entire heat exchange surface is 670W/m2/K, which is 49% higher than before optimization. Although the blade can increase the CHTC at the bottom, it also increases friction loss. The pressure drop between the inlet and outlet of the optimized structure increased by 2.3% compared with the original structure. In the case of a rotational speed of 3000rpm and a flow rate of 3 L/min, the comparison of specific data before and after optimization is shown in Table II. The rotor experimental platform shown in Fig. 2 uses eight heating pipes evenly distributed on the outer edge of the rotor, which is used to equivalent the actual rotor loss. In order to highlight the optimization of the cooling effect of the recirculating hollow-shaft, the rotor experimental platform was modeled and the temperature field simulation was performed using Workbench software. It can be seen from Fig. 17 that the rotor average temperature of the optimized structure is lower than that of the original structure, and the temperature difference between the two increases as the rotor loss increases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002850_s00170-020-05095-2-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002850_s00170-020-05095-2-Figure8-1.png",
        "caption": "Fig. 8 Structure and experimental arrangement",
        "texts": [
            " To verify the effects of the proposed optimization method of heat transfer and thermal deformation, a gear form grinding machine was optimized by Qin Chuan Machine Tool Group Co., Ltd. based on the proposed optimization method. All the materials that were used in the spindle, the contact pressure of each contact pair, and the surface roughness of each component were optimized according to Table 2. And the constrained mode of these three sets of bearings are that the middle bearings are fixed but other two sets of bearings can move freely along the axial direction. Figure 8 shows the structure of the gear form grinding machine. To measure the temperature and thermal elongation of the spindle, three temperature sensors and an amesdial were arranged on the spindle and gear as shown in Fig. 8. In order to ensure the measurement accuracy, the amesdial was mounted by a magnetic stand, and the gear was fixed to ensure the stability of the magnetic stand. In order to ensure the stability of the test device under All bearings are freeAll bearings are fixed NODAL SOLUTION STEP=1 SUB =1 TIME=3600 UX (AVG) RSYS=0 DMX =.304E-04 SMN =-.227E-04 SMX =.252E-04 -.227E-04 -.120E-04 -.137E-05 .927E-05 .119E-04 -.173E-04 -.669E-05 .395E-05 .146E-04 .252E-04 -.923E-04 -.646E-04 -.369E-04 -.917E-05 .185E-04 -"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000638_j.proeng.2015.08.069-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000638_j.proeng.2015.08.069-Figure2-1.png",
        "caption": "Fig. 2. Light propagation through a photoelastic model.",
        "texts": [
            "The temperature of the oven is increased with a speed of five degrees per hour up to the stress freezing temperature (125 degrees C.) The temperature is maintained during 10 hours to allow equilibrium. The temperature is then lowered slowly (5 degrees/hour) to room temperature; the load should be maintained during the entire test. The model is then mechanically sliced to allow fringe analysis on a regular polariscope. We used plane polarized light to obtain the isochromatic and the isoclinic fringes in order to determine the stress values and the stress directions. The light intensity obtained on the analyzer (Fig. 2) is given by the following relation (equation 1) [9] : 2sin2sin 22I (1) The terms 2sin2 and 2sin2 give, respectively, the principal stresses directions and the values of their difference with the following relation (equation 2): e fN\u03c3\u03c3 21 (2) Where N is the fringe order obtained experimentally from the isochromatic fringe pattern, e is the slice thickness and f is the fringe constant which depends on the light wavelength used and the optical constant C of the model material ( ). The value of the fringe constant f is determined experimentally with a disc of the same material as the model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002959_aeat-05-2019-0094-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002959_aeat-05-2019-0094-Figure2-1.png",
        "caption": "Figure 2 A double-helical gear pair system",
        "texts": [
            " Kinetic energy is expressed as follows: Td \u00bc 1 2 m h _V d 2 1 _Wd 2 1 _Ud 2 i 1 1 2 IdD _Bd 2 1 _Cd 2 h i 1 1 2 IdP\u00f0X1 _ad\u00de\u00f0 _BdCd _CdBd\u00de 1 1 2 IdP\u00f0X1 _ad\u00de2 1med X1 _ad\u00f0 \u00de _V dsin Xt1ad 1 c d\u00f0 \u00de h i med X1 _ad\u00f0 \u00de _Wdcos Xt1ad 1 c d\u00f0 \u00de h i (1) m is the mass, IdD and IdP are transverse mass moments of inertia and polar mass moment of inertia, respectively, X is spin speed of the shaft, ed is the magnitude of mass eccentricity and c d is phase angle relative to the Y axis. After substituting the kinetic energy (1) into Lagrange\u2019s equation, the equation of motion is obtained as follows: Md\u00bd f\u20acqg1X Gd\u00bd f _qg \u00bc fFdg (2) where [Md] is the mass matrix, [Gd] is gyroscopic effect matrix, {Fd} is the force vector due to disk eccentricity and {qd} is the displacement vector. As shown in Figure 2, the gears were devised as two rigid disks connected by a spring and a damper along their pressure line. In the model, the meshing stiffness coefficient km is considered as a constant, and the meshing damping coefficient cm is assumed to be zero. Therefore, the meshing gear force along the pressure line is as follows: Fh \u00bc kmd (3) The meshing gear displacement along the pressure line is as follows: d \u00bc Va Vb\u00f0 \u00desinc ij 1 Wa Wb\u00f0 \u00decosc ij cosb ij 1 Ua Ub\u00f0 \u00de1 raBa 1 rbBb\u00f0 \u00desinc ij sinb ij 1 raaa 1 rbab\u00f0 \u00de cosb ij 1 raCa 1 rbCb\u00f0 \u00decosc ij sinb ij (4) where ra and rb represent the radii of the driving gear and driven gear, respectively; c ij represent the pressure angle; Helix angle b ij is defined as: b ij > 0 if gear i has left hand teeth \u00bc 0 if gear pair is a spur gear < 0 if gear i has right hand teeth 8>< >: (5) The equation ofmotion for the gear pair is as follows: Mg1\u00bd 0 0 Mg2\u00bd \" # fqg1X1 Gg1\u00bd 0 0 Nt1 Nt2 Gg2\u00bd 2 64 3 75f _qg1 km Sh\u00bd fqg \u00bc fFdg (6) where [Mg1] and [Mg2] are the mass matrices of the driving gear and driven gear, respectively; [Gg1] and [Gg2] are the gyroscopic effect matrices of the driving gear and driven gear, respectively; {Fd} is the force vector due to disk eccentricity; [Sh] is the meshing effect matrix of the gear pair; Nt1 and Nt2 are the number of teeth of the driving gear and driven gear, respectively; and X1 is the spin speed of driving shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000233_s10846-019-01130-x-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000233_s10846-019-01130-x-Figure5-1.png",
        "caption": "Fig. 5 a Polaris Ranger XP and b driving course",
        "texts": [
            " Aligned with the front-side view camera, they provided DRC-Hubo+ with 180\u25e6 camera-view of driving course. 4. x-IMU with 3-axis accelerometer, 3-axis magnetometer and 3-axis gyro. Data rates up to 512 Hz Section 5 presents how the collected data (from the sensor head) be processed to provide more intuitive perception and to enable DRC-Hubo+ to drive the vehicle in a fast and safe manner. The first task in DRC-Finals was vehicle driving. In the task, contestant robots were asked to drive the utility vehicle through the outdoor driving course. Figure 5 demonstrates the ground-vehicle (Polaris Ranger XP 900) and the built driving course2 which were used in the competition. For the first task, the robot began in the vehicle, drove through the course, and crossed the finish line. The task was considered complete when both rear wheels of the vehicle have crossed the finish line. As described in Section 1, the authors developed a new perception data processing system (driving information module of control system architecture in Section 3) which is optimized for driving task above"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000970_j.electacta.2014.02.088-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000970_j.electacta.2014.02.088-Figure1-1.png",
        "caption": "Fig. 1. Experimental setup to measure the response in the Pruss",
        "texts": [
            " Each H2O2 test solution was prepared immediately before immersing the optical fiber with the multilayered sensing structure in it regardless of the temperature of the test. 2.3. Test for temperature response The systematic testing of the sensor involving both changes in temperature and concentration was experimentally impractical at this stage of development, but prior to the temperature response experiments, the fiber response to concentration was checked and a log-linear behavior was obtained consistent with the results reported in [19]. The experimental setup to test the response to temperature is shown in Fig. 1. The white light source (HL-2000-FHSA, OceanOptics, Dunedin, FL) was connected to the arm 1 of a bifurcated optical fiber (BIF200-UV-VIS, OceanOptics). Light was carried to the common arm that is connected to the sensing probe in one of the optical fibers SFS105/125Y. The reflected light was carried back through the common arm and into arm 2, where the spectrum was measured using an optical fiber spectrometer (USB2000, OceanOptics). Spectra from the sensing probe were recorded at a sampling rate of 0",
            " The temperature at which the tests were performed was controlled by placing the sensing probe inside an environmental chamber (TestEquity, Model 123H, Moorpark, CA). High density polyethylene (HDPE) containers were filled with either 3.0 mL of ABS at a pH 4.0 or 3.0 mL of ascorbic acid solution at pH 4.0 and placed inside the environmental chamber. The temperature of the solutions was monitored with a K-type thermocouple in a control bottle with 3.0 mL of ABS. Measurements were made at the time of the immersion of the probe. The uncertainty in these measurements is \u00b10.1 \u25e6C. Fig. 1 also shows how the sensing probe is held in 418 J.F. Botero-Cadavid et al. / Electrochimica Acta 129 (2014) 416\u2013424 ian blu a p t b H s o i l c s a i w l s t t T p 2 p c s t 5 t t 2 i o r t S a i vertical position, and how the immersion in the liquid reagents is erformed by raising a lab scissor jack stage. The experiment consisted of a series of cycles with an oxidaion stage and a reduction stage. During each oxidation stage PW ecomes PB after immersion of the sensor in solutions of 100 M 2O2. The reduction stage was induced by immersing the senor in ascorbic acid, which reduces PB back to PW"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002989_s12541-020-00344-6-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002989_s12541-020-00344-6-Figure1-1.png",
        "caption": "Fig. 1 Rok-2\u2019s assembly model (top left), gear part (top right), RoK-2 with ski suit (bottom left), and RoK-2\u2019s 3-D upper body assembly model (bottom right)",
        "texts": [
            " The ski robot RoK-2 is a humanoid robot and its height, including head, is about 132 cm and weight is 26 kg. The upper body of RoK 2 has a torso and two arms. The arms were made using 3-D printers to ensure they were light in weight. Each arm has three joints, including an elbow. The robot also has hands that can hold the ski poles. RoK-2 has two legs, each of which has three DoF at the hip, one at the knee, and two at the ankle. It was considered that the DoF configuration of the lower body could fully implement the human skiing motion. The overall appearance of the robot is as shown top left of Fig.\u00a01, and the specifications are in Table\u00a01. Skiing is a motion that creates a heavy load on the lower body in comparison with normal walking\u00a0[6], so sport skiers exercise lower body to raise strength in lower body for improved performance. Similarly, humanoid robots need additional strength in the whole lower body if they are to ski as well as walk. We have designed for maximum lower body strength and minimum weight. As in the top right of Fig.\u00a01, the ski robot\u2019s legs were operated with a dual mode using two motors for each joint, and a 1:3 reduction gear was added to exert greater torque to the hip roll, knee, and 1 3 ankle pitch joints. This addition was important for the ski motion to action extension and flexion in the legs and rotate the legs from the center of the hip roll. To lighten the lower body of the ski robot, the thigh and calf links were designed with light and durable carbon pipe. Ski robots need to be protected against the cold and humid snow. Instead, of casing the robot for waterproofing, we made a ski suit, as shown at the bottom left at Fig.\u00a01. The ski suit was waterproof for the robot while facilitating removal for inspection and maintenance. To prevent shorting of electronic components by snow or rain that might permeate into the ski suit, various electronic parts such as the control PC, the image processing PC, the electronic board, the inertial measurement unit (IMU) sensor, and the battery were mounted inside the robot body, which was then sealed. We checked the recommended operating temperatures of the electronic components needed for the robot, such as actuators, controllers, and sensors and selected the parts that continued to function at temperatures below 0 \u25e6C "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000760_icacci.2014.6968225-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000760_icacci.2014.6968225-Figure1-1.png",
        "caption": "Fig. 1: Parallel misalignment between two shafts.",
        "texts": [
            " There are two vari eties of misalignments: parallel misalignment and angular misalignment. The user should additionally distinguish between the two varieties of misalignment and imbalance once acting the machinery fault identification. The vibra tion level owing to imbalance can increase in proportion to the square of the speed whereas vibrations owing to the misalignment do not change. A. Diagnostics of Misalignment 1) Parallel misalignment: In this case the misaligned centerlines are parallel however not coincident (Figure 1). (If the machine speed will be varied, the vibration as a result of imbalance can vary with the square of the speed). Misalignment, resulting from parallel shifting, has simi lar symptoms as associate angular misalignment, however large vibration is within the radial direction. They're around out-of part, i.e. shifted by 180 0 over the coupling. The 2X part is commonly larger than the IX, however its size relative to the IX is commonly determined by the sort and construction of the coupling [13] (see Figure 3 below)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002840_012064-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002840_012064-Figure1-1.png",
        "caption": "Fig. 1. The design model: 1 \u2013 entrance surface; 2 \u2013 side surface;3 \u2013 bottom surface; 4 \u2013 symmetry surface; 5 \u2013 top surface; 6 \u2013 fused wire; 7 \u2013 weld bead; 8 \u2013exit surface",
        "texts": [
            "However, not enough literature is devoted to the CMT-pulsed welding process - a combination of CMT and pulsed arc, especially for welding aluminum alloys and galvanized steel. This technology has been known and used not so long ago, since 2004, so its comprehensive study, including process modeling, is an important and urgent task. In a first approximation, the mathematical model is based on the solution of the heat problem in a three-dimensional formulation in the first approximation. Due to symmetry, it is sufficient to include only half of the entire object in the computational domain for the description. The geometry of the computational domain is shown in Fig. 1. The following simplifications were made when modeling the surfacing process. Firstly, the entire volume of the metal under study is in a solid state, secondly, the heating source is written in the form of a Gaussian distribution, thirdly, the geometric parameters of the bead are associated with the parameters of the CMT process. The mathematical model is based on the solution of the differential heat equation (energy transfer) in a three-dimensional formulation ( ) (1) where T is the temperature, \u2013 coefficient of thermal diffusivity, V \u2013 medium velocity (matches surfacing speed and wire feed speed), \u03c1 \u2013 density, P \u2013 total power input to the product [W/m 3 ]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure18-1.png",
        "caption": "Figure 18. Airfoil (wing cross section) forces and terminology. (a) Forces on an airfoil moving at some angle of attack, \u03b1, relative to its direction of motion. D: drag; L: lift; R: resultant force. (b) A cambered airfoil of chord c with maximum height x of the mean camber line.",
        "texts": [
            " For the cost of overcoming a unit of drag force, the wing generates many units of lift force. Not only do all flying animals used wings to generate lift, but the propelling appendages of most macroscopic swimmers\u2014sea lion and sea turtle flippers, whale and tuna tails, blue crab hind legs\u2014function as underwater wings and generate lift. By convention, lift is defined as a force perpendicular to the direction of the wing\u2019s movement through the fluid. Drag is defined as parallel to the motion, so lift and drag are perpendicular (figure 18). Lift and drag can be added vectorially to give the resultant force, R. Biological wings are generally cambered (convex upward), and the degree of camber is given by x/c (usually expressed as a percentage) as in figure 18. Wings are relatively large structures that operate in fast flows so they are often streamlined; the sharp trailing edge not only helps reduce drag, but aids lift production as well, as described below. The amount of lift produced by a wing can be changed by changing the angle of attack, \u03b1 (figure 18). The angle of attack is the angle between the wing\u2019s chord line and the direction of movement. Within limits, the greater the angle of attack, the greater the lift. At the risk of somewhat oversimplifying, we can think of the wing\u2019s trailing edge directing air downward as it streams off the back edge of the wing (called \u2018downwash\u2019), and the downward momentum imparted to the air leads to an equal and opposite upward reaction on the wing, i.e., lift. Thus, increasing \u03b1 tilts the trailing edge down and increases lift, and giving the wing camber also tilts the trailing edge downward and increases lift"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000004_sice.2019.8859803-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000004_sice.2019.8859803-Figure1-1.png",
        "caption": "Fig. 1. Overview of the wireless pneumatic artificial muscle driver.",
        "texts": [
            " Adjusting the delay and active duration of the PAM leads to the optimal walking support taking individual differences of walking into account. Therefore, we developed wireless PAM driver that can adjust delay and active duration of PAM contraction by an intuitive operation by a smartphone. The purpose of this study is to perform the preliminary evaluation of the effect of the contraction parameters on kinematic profiles of the hip and knee joint during walking for achieving the optimal swing supprt. 2. SYSTEM OVERVIEW Figure 1 shows an overview of the wireless PAM driver developed in this study. Wireless sensor and display modules developed by Tada [13] are employed for realizing wireless and compact set up. a. Shoe part Pressure sensor (FSR 400 short; Interlink Electronics, United States) is used to detect the foot contact over the ground. This sensor is connected to an AD converter module (28.0\u00d739.5\u00d713.2 mm, 18g). The output of this sensor is sampled by this module through a voltage divider. The sampled data is sent to a smartphone as an OSC message via Wi-Fi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003583_biorob49111.2020.9224389-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003583_biorob49111.2020.9224389-Figure4-1.png",
        "caption": "Fig. 4. To fabricate a soft immobilization unit, a thin plastic film is perforated, folded in the (b) Miura fold pattern and then (a) sealed in an airtight skin with an air hose connector and a connection strap at both ends. (c) Evacuating actuator leads to a contraction predefined by the fold pattern of the skeleton and if blocked a force is exerted. (d) A shoulder unit at ambient air pressure is shown. (e) shows the unit fully contracted.",
        "texts": [
            " In this study, we focused on a prototype for breast cancer treatment. The shoulder and abdominal region were chosen as the location for the immobilization units to provide access to the breast region while preventing slouching motion of the patient in sitting position. A new method of soft immobilization was developed which meets the requirements of the system. The device was constructed from several individual units for different areas of the patient\u2019s body. Each unit is a fluid driven origamiinspired artificial muscle (FOAM) (Fig.4a) [9] with a Miura fold (Fig.4b) [22] actuated by negative relative air pressure (Fig.4a). This type of actuator has been shown to have good durability, and no performance decrease over 30000 usage cycles [9]. The Miura fold is chosen due to its in-plane contraction, ability to easily bend around an object, e.g., a shoulder, in the direction of contraction, and its robustness. Although typically used for actuation, by fixing the ends of the actuator to a base frame, thus restricting the lateral contraction, force is exerted normal to the actuator (Fig.4c). Each unit was fabricated by folding a 0.18 mm thick Polyester film in the Miura fold pattern to form the skeleton of the actuator. The plastic film was perforated with the Miura fold pattern by a laser cutter (Universal Laser Systems, Inc.) after which it could then be folded by hand. In the next step the skin of the actuator was prepared by attaching a plastic air hose coupler. The skeleton was then sealed within a coated, airtight nylon-fabric sheet by an impulse heat sealer (American International Electric, Co.). The unit was completed by gluing connectors to both ends of the skin. To characterize a single, soft immobilization unit, the contraction ratio of the unit was first determined by evacuating the unit to (\u221299\u00b11) kPa (Fig.4d and e). A contraction ratio of approximately 75 % was achieved. In a subsequent experiment, the holding force in relation to the applied pressure was investigated (Fig.5). A phantom representing the human 983 Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 19:14:19 UTC from IEEE Xplore. Restrictions apply. shoulder region was made of sheet wood (half circle with radius of (20.0\u00b10.1) cm) and attached to the load cell of an Instron 5944 (Instron, U.S.A) universal testing machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure9-1.png",
        "caption": "Fig. 9. The kinematic path of the output variable in plane-symmetric Mode II.",
        "texts": [
            " The output variable \u03b83 = \u03b87 can be obtained according to the loop closure equation ( Appendix A.3 ) as follows: cos \u03b87 = \u2212\u22122 + 6 cos \u03b85 + cos ( \u03b85 \u2212 2 \u03b88 ) + 2 cos 2 \u03b88 + cos ( \u03b85 + 2 \u03b88 ) 6 \u2212 2 cos \u03b85 + cos ( \u03b85 \u2212 2 \u03b88 ) + 2 cos 2 \u03b88 + cos ( \u03b85 + 2 \u03b88 ) (25) When l = 100 mm, l 1 = 141.42 mm, the range of \u03b8A 1 , \u03b8A 2 , \u03b8A 3 , and \u03b8A 4 are (0, \u03c0 ). Thus, the scope of \u03b82 , \u03b86 , \u03b84 , \u03b88 is (- \u03c0 , \u03c0 ), which is obtained according to Eq. (2) , and the scope of \u03b81 , \u03b83 , \u03b85 , \u03b87 can be obtained as (- \u03c0 , \u03c0 ). Therefore, the kinematic path of the output variable is shown in Fig. 9 . According to Eq. (2) , when \u03b8A 1 = \u03b8A 2 = \u03b8A 3 = \u03b8A 4 = 20 \u00b0, \u03b82 = \u03b84 = \u03b86 = \u03b88 = 68.43 \u00b0; when \u03b8A 1 = \u03b8A 2 = \u03b8A 3 = \u03b8A 4 = 30 \u00b0, \u03b82 = \u03b84 = \u03b86 = \u03b88 = 35.26 \u00b0; and when \u03b8A 1 = \u03b8A 2 = \u03b8A 3 = \u03b8A 4 = 30 \u00b0, \u03b82 = \u03b84 = \u03b86 = \u03b88 = 10.48 \u00b0. The kinematic path curves of the output variables ( \u03b83 , \u03b87 ) can be plotted, as shown in Fig. 10 (a). The configurations (A, B, C, D, E) with \u03b81 = \u03b85 \u2265 0 and \u03b83 = \u03b87 \u2265 0 are called inward configurations, and the configurations (F, G, H, I, J) with \u03b81 = \u03b85 \u2264 0 and \u03b83 = \u03b87 \u2264 0 are called outward configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure31-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure31-1.png",
        "caption": "Fig. 31. A quadrangular prism deployable mechanism.",
        "texts": [
            " Plane-motion Mode III (corresponding to Fig. 14 ) are shown in Fig. 27 . Figs. 28 and 29 are the snapshots of plane-motion Modes IV (corresponding to Fig. 16 ) and V(corresponding to Fig. 18 ), respectively. Fig. 30 shows the snapshots of the special configurations(corresponding to Fig. 20 ) of this multiple- mode mechanism. Similar to Bricard-like mechanism [61] , the multiple-mode mechanism discussed above can also be used as a construction unit to construct deployable mechanisms. As shown in Fig. 31 , a quadrangular prism deployable mechanism composed of two identical multiple-mode mechanisms connected by eight spherical joints is constructed. During its movement, the two identical multiple-mode mechanisms are mirror-symmetrical. When the upper unit is turned inward, the quadrangular prism deployable mechanism can be folded from the configuration in Fig. 31 (a) to the configurations in Figs. 31 (d) and (e). When the upper unit is turned outward, the quadrangular prism deployable mechanism can be folded from the configuration in Fig. 31 (a) to the configurations in Figs. 31 (f) and (g). Different from the triangular prism deployable mechanism, the quadrangular prism deployable mechanism also can be folded into a plane as shown in Figs. 31 (b) and (c) due to the characteristics of the multiple-mode mechanism. Moreover, the multiple-mode mechanism can also be used as a sub-unit to construct deployable polyhedron mechanisms. The construction method, constraint conditions, and deployable mode will be introduced and analyzed in our future papers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000750_j.jweia.2014.11.004-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000750_j.jweia.2014.11.004-Figure7-1.png",
        "caption": "Fig. 7. Projectile specimens used in launching tests.",
        "texts": [
            " Based on a series of numerical simulations of the initial breakup of above ground magazine, Fan et al. (2014) plotted each occurring debris according to its length ratio as a probability density function (Fig. 6). As shown in the figure, the debris shapes of ratios of 1:1:1 to 3:1:1 dominate the debris cloud. Therefore, in order to validate the aerodynamic coefficients proposed by Chai et al. (2012) which were adopted by the DeThrow II programme, three specimens of a square block, a rectangular block and an irregular block with different shapes ratios within 1:1:1 to 3:1:1 (Fig. 7) were made and launched by using a high pressure gas launcher (Fig. 8). The general set up of the launching test is shown in Fig. 9. During the test, specimens were launched from the high pressure gas launcher with different launching velocities at a constant launching angle of 301 and almost zero angular velocity. In order to trace the flight trajectory of the block launched, two digital high speed high definition (HD) video cameras were employed to record the course of the flight (i.e. view of sight is perpendicular to the flight trajectory)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001663_1.4024235-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001663_1.4024235-Figure3-1.png",
        "caption": "Fig. 3 An example of singular configuration when det JP is null: points U1, U2, S1, and S2 belong to the same plane",
        "texts": [
            " For the kinematic constraints, our analysis calculates the determinants of Jacobian matrices Jh and JP and verifies if their values are null, which correspond to singular configurations. With respect to singularities of the first kind, the determinant of Jh becomes null when either h1\u00bc 0, h2\u00bc 0 or even both are null. Fortunately, due to constructive issues, these conditions are not feasible. However, depending on the parameter values, the parallel mechanism might reach singularities of the second kind, when the moving platform will perform motions in the vertical direction, but it will not be able to change its position in the other directions (Fig. 3). The parameters \u2018u1x and \u2018u2x must be negative, in order to avoid their occurrence. The authors, in a previous work [29], provided a geometrical interpretation of such singularities. Table 1 presents the passive and active joint limits for the parallel mechanism. Table 2 shows the geometric parameters related to the location of the spherical and universal joints. The volume of the available workspace is approximately 216 dm3. For the chosen parameters, the parallel mechanism will have not singular configurations within its workspace"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002141_j.proeng.2016.05.136-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002141_j.proeng.2016.05.136-Figure4-1.png",
        "caption": "Fig. 4. (a) Normal bearing (b) Inner race fault, (c) Outer race fault",
        "texts": [
            " The setup was run at the speed of 2710 rpm (45.16 Hz). The sampling frequency is 10000 samples per second. The sampling frequency was selected arbitrarily to some extent. Two sets of trial data was taken for each condition. The time domain vibration readings were exported to MATLAB for Time-frequency analysis. The sample vibration signals are shown in Fig. 3 and the signals are analyzed through time frequency toolbox, version 2 in Mat lab for STFT, ZAM. The various defects views of selected bearings are shown in Fig. 4. 4. Results and Discussions The signals under normal and damaged conditions such as inner race fault and outer race fault conditions are taken from the experimental setup as discussed in section 3.The time frequency energy distributions (spectrogram, ZAM) of the acceleration signals of vibration recorded from normal and faulty bearings are shown in Figs. 5-7. The results show the performance of spectrogram and ZAM representations in locating the energy distributions in time frequency domain. Different patterns are obtained for normal and different faulty conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002162_jjap.55.071701-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002162_jjap.55.071701-Figure5-1.png",
        "caption": "Fig. 5. (Color online) Electro-optic response of a cell with HAN configuration at different polarity of the in-plane applied electric dc field generated by IPS electrodes (cf. Fig. 2). The electrode pattern is shown on the top of the figure. The cell was filled with the binary BC=RL nematic mixture containing 90wt% bent core nematic and 10wt% calamitic nematic mixture MLC6608 (Merck). This binary nematic mixture possesses \u0394\u03b5 < 0. The crossed polarizars are oriented in such way that the optic axis at one side of the electrode is oriented along the polarizer (dark state) whereas at the other side of the electrode, the optic axis is approaching position 45\u00b0. with respect to the polarizers\u2019 transmission direction (bright state).",
        "texts": [
            " The reorientation of Pflexo, however, resulted in switching of the liquid crystal molecules in the plane of the substrate (in-plane switching) generating thus a twist deformation across the liquid crystal layer and thus in-plane rotation of the sample optic axis at angle \u03b8, which is given by Eq. (2). Field-induced rotation of the sample optic axis gives rise to electro-optic response if the cell is inserted between crossed polarizers. The electro-optic response of the sample at different field polarity of the applied dc electric field is shown in Fig. 5. If the field-induced deviation of the optic axis in the regions between the electrodes \u03b8 is 22.5\u00b0, being clockwise and anti-clockwise, respectively, then orienting the crossed polarizers with respect to IPS electrodes at 22.5\u00b0, the region at one side of the electrode will be in bright state with light transmitted intensity given by Io;1 sin 2 2 8 \u00fe h i ; whereas at the other will be in dark state with light transmitted intensity corresponded to Io;1 sin 2 2 8 h i : As seen, the switching has a polar character"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001032_j.proeng.2014.09.147-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001032_j.proeng.2014.09.147-Figure3-1.png",
        "caption": "Fig. 3. Topography display of the SUS 304 stainless foil.",
        "texts": [],
        "surrounding_texts": [
            "Traditional cold rolling theories with an assumption that arc remains a circular with enlarged radius and one neutral point fail in foil rolling process due to significant elastic deformation of work roll. Fleck and Johnson (1987) suggests that the work rolls deform to a non-circular profile and a neutral region of no-slip exists in the roll bite for finite reductions of the strip. A new theory for cold foil rolling is presented by Fleck et al (1992) that plastic reduction takes place in two zones at entry and exit, which are separated by a neutral zone in which the rolls are compressed flat. Tribological analysis is very important in cold rolling process. Le and Sutcliffe (2003, 2006) applied a tribological model for mixed lubrication in rolling of thin strip and foil, and indicate the friction coefficient on the asperity contact is related to a theoretical oil film thickness and secondary-scale roll surface roughness. Sadowski and Stupkiewicz (2010) report the combined effect of friction and macroscopic deformation on asperity flattening. Roll roughness and lubricant viscosity influence the loads during cold rolling (Dick and Lenard 2005). Theories of hydrodynamic lubrication can be applied on smooth surfaces in cold rolling (Lugt et al, 1993; Wilson and Walwit, 1972). Batalha and Filho (2001) presented quantitative characterization of the surface topography of cold rolled sheets. For scale-down size effect, Vollertsen and Hu (2009) analyze tribological size effects in sheet metal forming measured by strip drawing test. Deng et al (2001) use an open and closed pockets theory analyze size effect on material surface deformation behaviour in micro forming process. Changing lubrication can alter the mode of deformation during forming processes and change the mechanical properties of the final product.\nIn this study, the micro rolling deformation characterization has been investigated in 2-Hi reversing micro rolling mill. The influences of the feature sizes on oil film thickness have been discussed in the high strength material of SUS 304 stainless steel with ranging thicknesses from 50 \u03bcm to 100 \u03bcm. The analysis of the evolution of surface roughness and oil film thickness has been conducted experimentally and analytically.\nNomenclature\nH initial foil thickness h foil thickness after rolling\noil film thickness in inlet zone oil film thickness in outlet zone\ncontact zone length , deformed roll radius\nr rolling reduction in thickness v rolling velocity\nroll surface velocity entry velocity of the foil average flow strength tensile stress pressure-viscosity coefficient dynamic viscosity\nThe research on multi-pass rolling deformation behaviour of SUS 304 stainless steel is carried out on the 2-Hi reversing micro rolling mill. The new micro rolling mill shown in Fig. 1 includes gear system, paralleled motor, gap adjustment, micro rollers, shaft connection and control system. The rollers are driven by two separated motors featured by adjustable and flexible rolling speed ratio. The roll and rolled foil are degreased with acetone before each pass. The initial thickness of SUS 304 stainless steel foil is 100 \u03bcm and the foil width is 15 mm. The true stress-true strain curve of the foil obtained in tension test is 1250 MPa. Fig. 2 shows schematically the contact with areas of contact between the roll and strip and areas separated by an oil film.",
            "Table 1 lists parameters used in experimental and analytical models. The annealed austenitic stainless steel of 100 \u03bcm thickness was prepared as rolled materials, and the 2-Hi desktop-scale reversing mill was used for multipass rolling. The rolled foil was collected, measured and compared after each pass. The pass reductions in foil thickness after each pass were in variation according to the rolling conditions.\nIn the case of foil rolling, deformation may be considered fairly homogeneous across the thickness of the rolled material. Note that in this model, the foil thickness is considered to be small compared to the length of the roll gap and its transverse width, and there is no stress variation in foil vertical direction. The stress only varies as the foil",
            "moves along the length of the roll gap. The contact pressure and roll shape for one pass with significant roll deformation are illustrated in Fig. 4 obtained from Fleck\u2019s model (Fleck et al, 1987 and 1992). The pressure was taken to be near Hertzian and the deviation of the roll profile from flat was incorporated by using a modified mattress model. The elastic and contained plastic compression of the foil can be ignored in the vertical direction because the foil is much thinner than the elastic deformation of the rolls. The sharp pressure peak is observed and the direction changes at the neutral section where the pressure peaks.\nA characteristic feature of the pressure distribution for foil is the sharp pressure peak which occurs just beyond the end of the neutral zone (D position). The pressure peak causes the rolls to deform locally to a concave shape. The pressure increases steadily through the no-slip zone and the rolled material is in a state of contained plastic flow. The foil in the central flattened region (CD zone) goes from a contained plastic zone to an elastic no-slip zone characterized by maximum in the pressure in central flattened region and then to contained elastic slip zone. BC zone and DE zone are featured plastic deformation zones in the foil rolling.\nIn order to investigate the oil film thickness in lubricant rolling, the oil drop method was applied. The test involved dropping a known quality of oil on the foil surface at the beginning of rolling, rolling the strip in the rolling mill, taking a digital photograph of mark left on the foil after rolling. The drop volume divided by the surface area of the mark gives the finial thickness of the oil film between the roll and the foil in the bite. The lubricant film thickness at the end of the inlet is determined by the lubricant rheological properties, the rolling speed and the roll geometry (Lenard, 2007). Wilson and Walowit (1972) derive an expression for the oil film thickness as below. Contact zone length can be obtained from the Fleck model and the ratio =\n/ of the smooth film thickness to the combined roll and initial strip roughness = + is used to characterize the lubrication regime.\n= ( ) ,\n[ ( )] 4T . (1)\nA film thickness at the outlet of the contact zone of\n= 4T . (2)\nThe calculated film thickness from mixed film model was compared with the measured film thickness shown in Fig. 5. Both of the oil film thicknesses in experiment and calculation decrease with increasing pass reduction in thickness. The calculated value is always smaller than the experiment measured value. The main reason is the film"
        ]
    },
    {
        "image_filename": "designv11_30_0002511_j.jart.2016.09.006-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002511_j.jart.2016.09.006-Figure6-1.png",
        "caption": "Fig. 6. Synoptic diagram of the \u201cball and beam\u201d system.",
        "texts": [
            " The ball and beam ystem is an unstable nonlinear system which is well known in utomation; therefore, it is regarded as a perfect bench test for he design of control laws for non-minimum phase nonlinear ystems. The ball and beam system is composed of a rigid bar carrying ball. The latter is characterized by its horizontal axis and its oment of inertia J. Its rotation angle \u03b8 compared to the horiontal one is controlled by an engine with direct current to which t applies a couple \u03c4. A ball is placed on the beam where it is ble to move with a certain freedom under the effect of gravity, s it is illustrated in Figure 6. The dynamic model governing the behavior of the ball and eam system in open loop can be expressed by the following quations (Hauser et al., 1992): i u w y \u03b7 Z b Constant 0.7143 ( Jb Rb 2 + Mb ) r\u0308b + MbG sin(\u03b8) \u2212 Mbrb\u03b8\u0307 2 = 0 (Mbr 2 b + J + Jb)\u03b8\u0308 + 2Mbrbr\u0307b\u03b8\u0307 + MbGrb cos(\u03b8) = \u03c4 (57) The numerical parameters of the ball and beam system are ecapitulated in Table 2. To define a new input u the system can e written in state-space form as (Hauser et al., 1992):\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u03071 x\u03072 x\u03073 x\u03074 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 x\u0307 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x2 Bb(x1x 2 4 \u2212 G sin(x3)) x4 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 f (x) + \u23a1 \u23a2\u23a2\u23a2\u23a3 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 g(x) u y = x1 (58) ith x = [ rb r\u0307b \u03b8 \u03b8\u0307 ]T , y = h(x) and Bb = b/(Jb/R2 b + Mb)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003890_icems.2013.6713239-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003890_icems.2013.6713239-Figure1-1.png",
        "caption": "Fig. 1. 3-Inertia Model.",
        "texts": [
            " The DE method proposed by Storn and Price proved to be a powerful global optimization technique. The DE makes it possible to give quick search of the required scaling factors and control gains and as a result to obtain the accurate control results. Finally, the effectiveness of the proposal method is confirmed by the computer simulations. II. 3-INERTIA VIBRATION SUPPRESSION CONTROL SYSTEM A. 3-Inertia Model The 3-inertia model, which consists of three rigid inertias with two torsional shafts, is shown in Fig. 1, where \u03c9M is the angular speed of the motor, \u03c9c is the angular speed of load 1, \u03c9L is the angular speed of load 2, Tin is the input torque, TL is the disturbance torque, JM is the motor inertia, Jc is the inertia of load 1, JL is the inertia of load 2, T1 is the torsional torque of shaft 1, T2 is the torsional torque of shaft 2, Ks1 is the stiffness of shaft1 and Ks2 is the stiffness of shaft2. In this research, we consider the current loop for the high speed torque control. And we have neglected the viscous frictions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002959_aeat-05-2019-0094-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002959_aeat-05-2019-0094-Figure3-1.png",
        "caption": "Figure 3 Shaft beam element and the coordinate system",
        "texts": [
            " Therefore, the meshing gear force along the pressure line is as follows: Fh \u00bc kmd (3) The meshing gear displacement along the pressure line is as follows: d \u00bc Va Vb\u00f0 \u00desinc ij 1 Wa Wb\u00f0 \u00decosc ij cosb ij 1 Ua Ub\u00f0 \u00de1 raBa 1 rbBb\u00f0 \u00desinc ij sinb ij 1 raaa 1 rbab\u00f0 \u00de cosb ij 1 raCa 1 rbCb\u00f0 \u00decosc ij sinb ij (4) where ra and rb represent the radii of the driving gear and driven gear, respectively; c ij represent the pressure angle; Helix angle b ij is defined as: b ij > 0 if gear i has left hand teeth \u00bc 0 if gear pair is a spur gear < 0 if gear i has right hand teeth 8>< >: (5) The equation ofmotion for the gear pair is as follows: Mg1\u00bd 0 0 Mg2\u00bd \" # fqg1X1 Gg1\u00bd 0 0 Nt1 Nt2 Gg2\u00bd 2 64 3 75f _qg1 km Sh\u00bd fqg \u00bc fFdg (6) where [Mg1] and [Mg2] are the mass matrices of the driving gear and driven gear, respectively; [Gg1] and [Gg2] are the gyroscopic effect matrices of the driving gear and driven gear, respectively; {Fd} is the force vector due to disk eccentricity; [Sh] is the meshing effect matrix of the gear pair; Nt1 and Nt2 are the number of teeth of the driving gear and driven gear, respectively; and X1 is the spin speed of driving shaft. This paper used the Timoshenko beam in developing the model. The effect of shear deformation and rotational inertia were considered, and the finite element method was adopted to derive the equation of motion of the shaft. Figure 3 shows that one node of the shaft has six degrees of freedom, consisting of one axial displacement, two lateral displacements and three torsional displacements. One shaft unit is composed of two nodes; therefore, one finite shaft element has 12 degrees of freedom. The kinetic and strain energy of the shaft unit are expressed as follows: Tr \u00bc 1 2 \u00f0 l 0 rA _V r 2 1 _Wr 2 1 _Ur 2h in o ds1 \u00f0 l 0 rIrD _Br 2 1 _Cr 2h in o ds 1 2 \u00f0 l 0 X1 _ar\u00f0 \u00derIrP _CrBr _BrCr ds1 1 2 \u00f0 l 0 X1 _ar\u00f0 \u00de2rIrPds (7) Double-helical geared rotor system Ying-Chung Chen, Tsung-Hsien Yang and Siu-Tong Choi Aircraft Engineering and Aerospace Technology Volume 92 \u00b7Number 4 \u00b7 2020 \u00b7 653\u2013662 where A is the cross-section area, r is the mass density of the shaft, IrD and IrP are the transverse moment of inertia and polar moment of inertia of the shaft, respectively, E is Young\u2019s modulus of the shaft, k is shear factor for the circular crosssection of the shaft andG is the shearmodulus of the shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.6-1.png",
        "caption": "Fig. 18.6 Hydrogen reference electrode",
        "texts": [
            " Nevertheless, even in those cases a finite charge-transfer resistance exists at such an interface and excess charge leaks across with the time constant given by the product of the double-layer capacitance and the charge-transfer resistance (\u03c4\u00bcRctCdl). Many analytes are aqueous solutions, thus it makes sense to use a reference electrode where hydrogen ions participate in the reaction. A standard hydrogen electrode (SHE) consists of a platinum foil surrounded by a platinum sponge called \u201cplatinum black\u201d having very large surface area (Fig. 18.6). The platinum assembly is positioned inside a glass tube with a built-in small chamber where a drop of mercury is placed to make electrical contact between the platinum and copper wires for the external electrical connection. The Pt Black is dipped in the analyte solution for the electrochemical reaction. The glass tube has an inlet for pumping in H2 gas, while the base is perforated for escaping the excess of hydrogen. When pure, dry hydrogen gas is passed through the inlet, it flows by the Pt Black sponge",
            " Since the e-nose and e-tongue sensing cells in an array are relatively slow to respond, produce intrinsic noise and have relatively low selectivity, the bionic methods of signal processing become more popular. They employ the adaptive and learning (trainable) neural-network software in the DSP and can yield quite impressive results by responding to dynamics of the changing outputs of the sensing array [111]. This neural processing has three advantages: faster speed response, improved signal-to-noise ratio, and better selectivity. Figure 18.6 illustrates an array of multiple CP sensors that are exposed to an odorant [112]. The dynamical approach utilizes time transients of the sensor responses without waiting for their settling on constant levels. Each sensor in the array is predominantly sensitive to a specific odorant and coupled to one or more inputs of the neural network that responds to the rates of the signal changes and their magnitudes (Fig. 18.34). Noisy and slow sensor responses are refined by the neural system formed as a multiple coupling of the excitatory (white circles) and inhibitory (black circles) dynamical units"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002882_s11581-020-03504-w-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002882_s11581-020-03504-w-Figure6-1.png",
        "caption": "Fig. 6 Model of double crystal bridge Fig. 8 Deformation of Ag-IPMC under 3 V voltage",
        "texts": [
            " Therefore, if the IPMC deformation parameters are directly substituted into the bimorphmodel, a large error will occur. To solve this problem, the academic community proposes an equivalent thermal deformation to deal with its deformation. The beam model Table 2 Comparison between simulation of displacement and actual displacement Moisture content (%) Calculation displacement (mm) Actual displacement (mm) Relative error (%) 0 17.58 13.79 27.42 8.07 18.57 17.32 6.67 13.43 18.84 18.18 3.61 composed of the bimorph is shown in Fig. 6. The piezoelectric coefficient equation can be obtained by calculation: s \u00bc 3L2 2H2 1\u00fe b\u00f0 \u00de 2b\u00fe 1\u00f0 \u00de ab3 \u00fe 3b2 \u00fe 3b\u00fe 1 d31E3 \u00f013\u00de where a is the ratio of the modulus of elasticity, b is the ratio of the membrane thickness, H is the thickness of the model, L is the length of the beam/mm, E3 is the electric field strength/ (Pc/N), d31 is the piezoelectric coefficient/(N/C), and s is the tip displacement of the beam/mm. Due to the continuous deformation of the piezoelectric coupling, the elastic layer can be ignored"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002956_s11831-020-09423-3-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002956_s11831-020-09423-3-Figure5-1.png",
        "caption": "Fig. 5 Block diagram of the 1-DOF system: a in vertical plane, b in horizontal plane [19, 30]",
        "texts": [],
        "surrounding_texts": [
            "The 2-DOF coupling control is realized when the two locking screws are removed, giving way for the beam of the TRMS to have unrestricted movements and rotations in all directions and planes (vertical and horizontal). The schematic is shown in Figs.\u00a04a and 6. The main goal of the attitude coupling control, as in the case of the 1-DOF, is to situate the beam dynamically in the horizontal direction by implementing control algorithms on the pitch and yaw axes. Once this position is achieved, the counterbalance provides the opposing force/torque about the beam\u2019s pivot joint. Abdulwahhab and Abbas proposed an optimization tuning approach based on a Fractional Order PID ( PI D ) control for the 2-DOF yaw and pitch rotations of the Twin Rotor Aerodynamical System (TRAS) [15]. Due to the inherent high order nonlinearities present in the TRAS system, it was first linearized in order to obtain a suitable operating point, and decoupled into 2 subsystems. A conventional (i.e. integer-order) PID controller is obtained when \u03bb = \u03bc = 1; when 1 3 \u03bb = 1, \u03bc = 0 we get a PI controller; and when \u03bb = 0, \u03bc = 1 we obtain a PD controller. Hence the PI, PD and PID controllers are special cases of the PI D controller, or they are special points on the PI D control plane. This approach can be extended to systems with packet dropouts and quantization, time delays, nonlinear fuzzy dynamic models, and to fractional order systems with complex and parameters. A hybrid intelligent controller, composed of a fuzzy PID controller with Real-valued GA (RGA) for fast attitudes control of the pitch and yaw (or azimuth) angles of the 2-DOF TRMS [19] was proposed. Kannan and Sheenu [52] proposed a root locus technique using pole placement approach to design and implement a PID controller for the 2-DOF TRMS. The pole assignment ensures that the desired performance specifications are met which is within the stability margins of the closed loop response. Hence with the help of the PID controllers, the tracking errors were significantly reduced, for both the linear and nonlinear consideration of the TRMS. Toha and Tokhi proposed a parametric technique using PSO-with Spread Factor (PSO-SF) to develop an inverse model and PID optimization as an intelligent approach to control the 2-DOF TRMS [10]. The augmented PID controller performed better using the feedforward and feedback inverse model by shortening the response and settling times of the conventional PID controller. The method has shown to be robust with good disturbance rejection capability. Ghellab et\u00a0al. proposed a Fuzzy Gain Scheduled PID controller (FGSPID) [32]. A 2-DOF simultaneous pitch and yaw motions attitudes control of the TRMS in hovering mode using fuzzy sliding controller (FSC) approach [27], with integral action (FISC) and a combined FSFISC) methods [17] were proposed. This is an integration of a fuzzy loci and SMC approaches. The uncertainties and external disturbances in the system was considered to be the cross-couplings existing between the vertical and horizontal directions. The merit to this approach is that the effect of chattering, which is the main concerns of SMC, is significantly reduced with this method. The use of the integral sliding mode control (FISC) arises from the fact that the TRMS model weight distribution is asymmetrical, with the main rotor heavier than the tail rotor. Hence the pitch angle would not be at the expected position when the system is not controlled. It is therefore not an easy task to control the vertical subsystem to attain zero position tracking error for the pitch angle, with a pure/conventional fuzzy logic sliding mode control (FSC). To cancel/remove the effects of the unequal weight distribution, an FISC is implemented. Da Silva et\u00a0al. proposed a loop-shaping sensitivity function procedure based on the H\u221e parameterization to control yaw and pitch angles of the 2-DOF TRMS [20]. It is based on similarity transformations for obtaining first-order system behavior through singular values sensitivity. The advantage of this method over other loop-shaping methods of earlier works is that for a process plant that is open loop stable, in the case of the TRMS, can be decoupled into 2 first-order systems where their open loops due to the open loop poles can be shaped directly. The limitation is that the decoupling of the system into 2 first-order systems can only be applied to stable open-loop systems and not to closed-loop and openloop unstable systems. A considerable number of control techniques substantially depend on the system states, e.g. inverse dynamics and feedback linearization. But for most physical systems whether linear or nonlinear, only inputs and outputs to the system can be measured. And we know that the system states play a pivotal role in system monitoring, fault detection and diagnosis and in achieving superior performance, hence the need for system states estimation. For the nonlinear TRMS, observer design based on NN approximation is more suited and recommended. A stable NN-based observer approach as a Nonlinear State Observer for the 2-DOF pitch and yaw motions was proposed in [16]. Sodhi and Kar implemented an Adaptive Backstepping Controller based on Lyapunov stability analysis using a State Observer Approach for the 2-DOF yaw and pitch attitudes angles of the TRMS [53]. The control objective is to precisely track the trajectory of the TRMS\u2019s beam so that it rests in desired horizontal and vertical positions at steady-state, under predetermined load conditions. The main control difficulty arises from the fact the TRMS: 1. Is highly nonlinear 2. Data obtained from the embedded incremental shaft encoder (with resolutions of 100 pulse/revolution) contains systemic measurement noise, due to modelling and parametric uncertainties and disturbances. A hybrid of the Luenberger type and Extended Kalman Filter (EKF) was applied to estimate all the system states. Lyapunov stability analysis ensures the closed loop system of the TRMS is stable asymptotically with the help of adaptive update law. With these, the pitch and yaw angle estimates are made, although the estimates for the yaw angle is completely inaccurate. Hence the hybrid controller is used to obtain a reliable estimate for the dual angles in question, in the presence of observation and systemic errors (i.e. the measurement noise presumed to be Gaussian with zero mean) and are cancelled-out by an adaptive fine-tuning process on the uncertain system parameters (in the covariance matrices of the Kalman Filter algorithm). Yang and Hsu proposed an Adaptive Backstepping Controller based on the Lyapunov stability analysis using an Adaptive Control Approach for the 2-DOF pitch and yaw attitudes (angles) control of the TRMS [25]. Kulkarni and Purwar [54] used an adaptive Composite Nonlinear 1 3 Feedback (CNF) controller for the 2-DOF TRMS with input saturation. The nonlinear gain was based on the adaptive law and the Lyapunov stability analysis for all nonlinear gain values of zero, negative or positive. For larger tracking errors, the controller performs well adequately."
        ]
    },
    {
        "image_filename": "designv11_30_0001488_amm.611.90-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001488_amm.611.90-Figure2-1.png",
        "caption": "Fig. 2 Crank Rocker Mechanism - velocity of the point A and velocity of the point B",
        "texts": [
            " 1 Velocities of the points A and B For the velocity of the point A For the velocity of the point B For the velocity vector vB31 of the member 3 vA31=vA32+vA21 vB=vA+vBA vB31=vB34+vB41 vA32=0 vB31=vA31+vBA31 vB34=0 vA31=vA21 vA31=vA21 vB31=vB41 Acceleration aB of the point B [6], [7]: BAAB aaa += (9) Where: aBt - known direction, aBn - can be obtained by Euclidean construction, aAt - is zero because ,0,0. 21212121 =\u22c5==\u21d2= AOakon\u0161t At \u03b1\u03b1\u03c9 (10) aAn \u2013 is normal component of the acceleration of the point A: , 21 2 AO v a A An = (11) Graphical solution of the above-stated vector equation is in Fig. 2, Fig. 3. Angular velocity \u03c941 and angular acceleration \u03b141 is then calculated according to: In Fig. 2 there is the model in graphic form with velocities of points A and B. In Fig. 3 a) there is the vector velocity of the point A in time t and the acceleration vector of the point A in time t, velocity and acceleration. Simulation of the crank rocker mechanism using MSC Adams/View The given crank rocker mechanism was modeled in MSC ADAMS/View and the initial parameters were provided [2],[3]. In the initial window of the program Adams we set data folder, name of the project, units and the working grid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001349_00032719.2014.900782-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001349_00032719.2014.900782-Figure2-1.png",
        "caption": "Figure 2. Effect of extraction time and salt addition at temperature of 50 C with single-wall carbon nanotubes=silica composite.",
        "texts": [
            " Single-Wall Carbon Nanotube/Silica Composite For single-wall carbon nanotubes=silica composite, the optimal extraction time was 60min. Temperature is an effective parameter on gas-aqueous and the gas-sorbent partitioning (Bagheri et al. 2011) Since toluene has low K value, at temperatures above 70 C, the rate of sorption was reduced. These results show that the best extraction temperature was 50 C. The results indicated that the best extraction conditions were obtained with 15% sodium chloride due to an increase in the ionic strength of aqueous solution (Bagheri et al. 2011) (Fig. 2). Polydimethylsiloxane For polydimethylsiloxane as a commercial sorbent, the best extraction condition was implemented in 30min and 30 C without sodium chloride. The results show that above 30min, the polydimethylsiloxane was blocked with analyte and aqueous solution so the extraction process was reversed. At high temperatures, Figure 1. Scanning electron microscopic image of single-wall carbon nanotube=silica composite. D ow nl oa de d by [ D ic le U ni ve rs ity ] at 0 8: 24 1 5 N ov em be r 20 14 toluene might easily separate from the aqueous solution, but above 40 C a decrease in the equilibrium of analyte was noted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000548_978-3-319-05431-5_3-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000548_978-3-319-05431-5_3-Figure8-1.png",
        "caption": "Fig. 8 Muscle recruitment for knee extension",
        "texts": [
            " The desired torque is imposed at the knee, the extension force is measured and the muscles are electrostimulated through a PI controller to reach the targeted torque. Meanwhile, extension movements (Green) are carried out by the motorized orthosis. The recruited muscles are the RF (Rectus Femoris), the VM and VL (Vastus Medialis and Lateralis). The obtained results have proved the possibility to follow an isotonic torque command during an isokinetic knee extension by closed loop electrical stimulation (Fig. 8). The MotionMaker has been used for the implementation of the strategy. This device allows flexion/extension movements for each leg. Each joint is equipped with a force sensor. The foot force is hence measured thanks to the Jacobian matrix J (Sciavicco and Siciliano 2000). Figure 9 illustrates how the closed loop force control works. First, a flexion or extension target force at the foot is set. This force corresponds to the desired torques at the joint levels. The 3 torque control loops work independently and the PI (Proportional + Integrator) current values are applied to the recruited muscles participating to the flexion or the extension of each corresponding joint (Metrailler 2005)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001352_chicc.2014.6896925-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001352_chicc.2014.6896925-Figure2-1.png",
        "caption": "Fig. 2: The coordinate system of thequadrotor UAV.",
        "texts": [
            " The attitude dynamics of the aircraft is under nonlinearities with uncertain inertial moments, aerodynamic coefficients and mass-eccentric constants The outline of this paper is listed as follows: In Section 2 and 3, the dynamic model of the quadrotor and the problem statement are presented. The design of the I&I based adaptive controller is given in Section 4, while the stability analysis is provided in Section 5. To demonstrate the performance of the adaptive flight control law, Section 6 presents the numerical simulation results. Finally, the conclusion can be found in Section 7. 2 Problem Statement The schematic of a quadrotor UAV is shown in Figure 2. The thrusts generated by the four rotors are denoted by fi (t) , i = 1, 2, 3, 4 respectively. Let I = {xI , yI , zI} represent a right hand inertia frame with zI being the vertical direction towards the sky. The body fixed frame, denoted by B = {xB , yB , zB}, is located at the center of gravity of the aircraft. Let \u03b7 (t) = [\u03c6 (t) , \u03b8 (t) , \u03c8 (t)] T \u2208 R 3 be the Euler angle of the UAV with respect to the frame I. The dynamic model of the quadrotor UAV under consideration in this paper is as followed: J1\u03c6\u0308+K1l\u03c6\u0307+ \u03b41 = lu1 J2\u03b8\u0308 +K2l\u03b8\u0307 + \u03b42 = lu2 (1) J3\u03c8\u0308 +K3\u03c8\u0307 + \u03b43 = cu3 where Ji \u2208 R for i = 1, 2, 3 are the moments of inertia of corresponding axes, Ki \u2208 R for i = 1, 2, 3 represents the aerodynamic damping coefficient, \u03b4i \u2208 R for i = 1, 2, 3 are parameters represented the eccentric influence caused by the unbalanced payload, and regarded as constants approximately when hovering within a small range of rotation angle, l denotes the distance between the epicenter of the UAV and the rotor axis, c \u2208 R represents a constant force-to-moment factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003052_s00773-020-00730-9-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003052_s00773-020-00730-9-Figure2-1.png",
        "caption": "Fig. 2 Horizontal plane path following",
        "texts": [],
        "surrounding_texts": [
            "Vol.:(0123456789)\nKeywords AUV\u00a0\u00b7 Path following\u00a0\u00b7 Dynamical sliding mode control\u00a0\u00b7 Vector field algorithm\u00a0\u00b7 ILOS\nPath following is required for AUV in many critical operations, such as intelligence collection, ocean investigation, pipeline inspection, ocean rescue, etc. [1]. To reduce the incidence of mechanical failure, manufacturing costs, weight and energy consumption, AUVs are often designed in the form of underactuated. Underactuated AUVs are more difficult to control because they lack the driving force on certain degrees of freedom [2]. In addition, the underwater motion of AUV is a six-degree-of-freedom movement in three-dimensional space, and the motion of each degree of freedom is nonlinear and coupled. Its precise motion model and hydrodynamic coefficient are difficult to obtain [3]. Moreover, AUVs work in challenging environments where disturbances, such as ocean currents, can significantly affect\ntheir maneuverability [4]. For these reasons, the path following control problem is widely studied in the world. Guidance system is a key technology for AUV to realize path following. At present, the more popular guidance principles are vector field (VF) guidance [5, 6] and line-of-sight (LOS) guidance [7, 8]. However, they are susceptible to unknown drift forces caused by external disturbances, such as ocean currents. When AUVs lack lateral and vertical driving forces and are affected by drift forces, the path following using LOS or VF guidance law will generate steady-state errors, so that AUVs cannot converge to the desired path eventually. At present, several methods are used to compensate for the drift forces. The most effective approach is to calculate directly the attack and sideslip angles using the measurement of velocities in surge, sway and heave [9]. Alternatively, control methods based on ocean current observers have been developed to compensate the drift forces [10\u201312]. However, the two methods are difficult to implement, because AUVs usually lack sensors due to the limitations of space or cost, and the measurement precision is difficult to be guaranteed. To make up for the shortcomings of the prior methods, integral line-of-sight (ILOS) guidance is an efficient way to compensate for the ocean current disturbances [13\u201315]. In [16], an adaptive controller has been developed for compensation * Xiaowei Wang wangxiaowei@hrbeu.edu.cn 1 College of\u00a0Automation, Harbin Engineering University, Nangang District, Harbin\u00a0150001, China\n2 College of\u00a0Mechanical Engineering, Jiujiang Vocational and\u00a0Technical College, Lianxi District, Jiujiang\u00a0332007, China",
            "1 3\nof the drift forces. The dynamic controller is also studied extensively. The PID controller was firstly applied to tackle the control for underactuated AUV [17]. Then, the controllers based on feedback linearizing control techniques [6, 10], back-stepping method [18, 19] and Lyapunov direct method [20] have been developed. Many researchers use adaptive control [21, 22], sliding mode control (SMC) [23\u201325], gain scheduling control [26], and robustness control [27, 28] to design dynamic controller, which can improve the robustness of the control scheme. According to the abovementioned papers, only the 2D path following problem or straight path following problem is solved. To complete different control tasks, the controller should be designed to make the AUV simultaneously track the straight and curve paths in 3D space.\nThe 3D path following control of underactuated AUV with ocean current disturbances and model uncertainties is studied in this paper. An integral vector field (IVF) algorithm and the virtual target with adaptive law are designed which can compensate for the drift forces and avoid singularities, respectively. The relative velocities between AUV and fluid are applied to design the dynamics controller, which can reduce the energy consumption. The dynamic controller is designed based on back-stepping method and adaptive dynamical sliding mode control (BADSMC) technology. Finally, the stability of the whole control system is analyzed by Lyapunov theory and nonlinear cascade system theory.\nIn remainder of this paper, the control task of 3D path following will be analyzed in Sect.\u00a02. In Sect.\u00a03, the 3D path following controller of underactuated AUV will be designed. In Sect.\u00a04, the stability of the system will be analyzed. In Sects.\u00a05\u20136, the simulation results and conclusions will be presented.\nIn this paper, the underactuated AUV has a propeller in the stern to realize the surge speed control. In addition, a pair of horizontal rudder is responsible for the control of pitch and a pair of vertical rudder is responsible for the control of yaw. The AUV has no driving forces in other degrees of freedom. The motion model of AUV in 3D are established in a moving coordinate system {B} \u2236 O \u2212 xyz and a fixed coordinate system {I} \u2236 E \u2212 , as shown in Figs.\u00a01, 2, 3. The center of buoyancy (CB) of AUV is set to the origin of {B} coordinate system; the coordinate of center of gravity (CG) in the {B} frame is [0 0 zg].\nAssumption 1. The linear velocities of ocean current in {I} coordinate system are defined as vI c = [ v v v ]T and\nsatisfies \u221a\nv2 \ud835\udf09 + v2 \ud835\udf02 + v2 \ud835\udf01 < Vmax , where 0 < Vmax . The angular\nvelocities of ocean current are zero. The kinematics and dynamics models of AUV can be simplified as the following [1] (1)\u0307 = J( )vr + [ v\ud835\udf09 v\ud835\udf02 v\ud835\udf01 0 0 0 ]T ,",
            "1 3\nwhere = [ ]T , the , , are the coordinates of CB defined in {I} frame, the , , are the orientation of AUV defined in {I} frame, \u03d5 is the roll angle, \u03b8 is the pitch angle, \u03c8 is the yaw angle. J( ) represents the rotation transformation from {B} to {I},\nvr = vB \u2212 vB c = [ ur vr wr p q r ]T is the relative velocity between AUV and fluid; ur, vr,wr are the surge, sway and heave relative velocities, respectively. vB = [ u v w p q r\n]T is the velocity of AUV defined in {B} frame; u, v,w are surge, sway and heave velocities; p, q, r represent the roll, pitch and yaw rates. vB c = [ uc vc wc 0 0 0\n]T is the velocity of ocean current defined in {B}, where [ uc vc wc ]T = JT 1 ( )vI c . The system matrices satisfy the properties M = MT , C ( vr ) = \u2212C ( vr )T,D ( vr ) > 0 . The restoring forces and\nmoments are defined as g( ) = [ 0 0 0 KHS MHS 0 ]T , where KHS = \u2212 zgGcos\u03b8sin\u03d5, MHS = \u2212 zgGsin\u03b8, G represents\n(2)Mv\u0307r + C ( vr ) vr + D ( vr ) vr + g( ) = + d,\nJ( ) = [ J1( ) 03\u00d73 03\u00d73 J2( ) ] ,\nJ1( ) = \u23a1 \u23a2\u23a2\u23a3 cos cos cos sin sin \u2212 sin cos cos sin cos + sin sin sin cos sin sin sin + cos cos sin sin cos \u2212 cos sin \u2212 sin cos sin cos cos \u23a4 \u23a5\u23a5\u23a6 ,\nJ2( ) = \u23a1\u23a2\u23a2\u23a3 1 sin tan cos tan 0 cos \u2212 sin 0 sin \u2215 cos cos \u2215 cos \u23a4\u23a5\u23a5\u23a6 .\nthe gravity of the AUV. = [ X 0 0 0 M N ]T represents the control inputs; X\u03c4 is the propeller thrust, M\u03c4 is the pitch moment, and N\u03c4 is the yaw moment. Vector d describes the model uncertainties. To make it convenient for the dynamic controller design, dynamic Eq.\u00a02 can be expanded as following\nwhere\nm is the mass of the AUV. Ixx, Iyy, Izz denote the moment of inertia. X{\u2219}, Y{\u2219}, Z{\u2219},K{\u2219},M{\u2219},N{\u2219} are hydrodynamic parameters. di(i = u, v,w, p, q, r) are model uncertainties.\nTo avoid the singularity problem of curve path following,\nSerret\u2013Frenet coordinate system {SF} \u2236 P \u2212 xsf ysf zsf is introduced as virtual target. The {SF} frame is a right-handed coordinate system, and the longitudinal axis represents the direction of its motion which tangent to the path, see Figs.\u00a01, 2, 3. Since the roll angle of AUV studied in this paper is very small, if the roll is ignored, the objectives in path following can be expressed as:\n(3) \u23a7 \u23aa\u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23aa\u23a9 u\u0307r = Fuur + FXX\ud835\udf0f + du, v\u0307r = Fvvr + dv, w\u0307r = Fwwr + dw, p\u0307 = Fpp + dp, q\u0307 = Fqq + FMMHS + FMM\ud835\udf0f + dq, r\u0307 = Frr + FNN\ud835\udf0f + dr,\nFu = Xu|u|||ur|| m \u2212 Xu\u0307 ,FX = 1 m \u2212 Xu\u0307 ,Fv = Yv|v|||vr|| m \u2212 Yv\u0307 , Fw = Zw|w|||wr\n|| m \u2212 Zw\u0307 ,Fp = Kp|p||p| Ixx \u2212 Kp\u0307 ,Fq = Mq|q||q| +Muqur Iyy \u2212Mq\u0307 ,\nFM = 1\nIyy \u2212Mq\u0307\n,Fr = Nr|r||r| + Nurur\nIzz \u2212 Nr\u0307\n,FN = 1\nIzz \u2212 Nr\u0307\n."
        ]
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.21-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.21-1.png",
        "caption": "Figure 2.21 Sintered EEG electrodes. (A) Ag AgCl sintered fixed wired electrode without housing. (B) Ag AgCl disc with housing (plastic cover); the disc face is recessed to apply the electrolytic gel. (C) Top view of Ag AgCl discs with and without a central hole.",
        "texts": [
            " In this process, a homogeneous mixture of fine powders of Ag and AgCl is compressed by a special technique, without the use of binders. As a result, these electrodes are homogeneous across the thickness, do not bend, have high mechanical strength and do not require rechloriding. They have wide contact area and present as further characteristics, very low offset voltage, polarization, rate of drift and noise level and are less susceptible to artifacts than the conventional electrodes (Tallgren, Vanhatalo, Kailaa, & Voipio, 2005). Figure 2.21 shows some examples of sintered EEG electrodes. Figure 2.21A shows sintered fixed wired electrode without housing and Figure 2.22B shows the electrode with plastic cover. Sintered discs are manufactured with and without a central hole (Figure 2.21C), in different disc diameters (4 12 mm) to suit to infant and adult use; they are coupled to the scalp skin with electrolytic gel and are fixed using collodion like the conventional electrodes. They are widely used with caps for brain mapping and long-term brain activity monitoring. The noninvasive EEG recording routines include some types of nonscalp electrodes, like the earlobe clips, used as reference, and nasopharyngeal electrodes, used in investigation of the origin from specific epilepsy activities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000743_s00542-013-1844-6-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000743_s00542-013-1844-6-Figure2-1.png",
        "caption": "Fig. 2 Finite element model and pressure distribution of the FDBs",
        "texts": [
            " Fx \u00bc mumex2cos\u00f0xt\u00de; \u00f016\u00de Fy \u00bc mumex2 sin\u00f0xt\u00de; \u00f017\u00de where mum is the unbalanced mass, e is the distance of the unbalanced mass from the mass center, and x is rotating speed. NRRO is generated due to various sources such as manufacturing errors or instability of FDBs and external shocks in real operation. This research estimated NRRO due to half-speed whirl by applying the swept sine excitation in addition to the centrifugal force as follows: F \u00bc p0sin a 2 t2 \u00fe xst ; \u00f018\u00de a \u00bc xe xs T ; \u00f019\u00de where a, T , xs and xe are sweep rate, sweep period, starting and ending frequencies. Figure 2 shows a finite element model of the FDBs of a 2.500 HDD with a rotating speed of 5,400 rpm and calculated pressure distribution of the FDBs. This finite element model consists of two grooved journal bearings, four plain journal bearings, two grooved thrust bearings, and one plain thrust bearing. Fluid film was discretized by 7,240 isoparametric bilinear elements with four nodes. The Reynolds boundary condition was applied to guarantee continuity of pressure and pressure gradient. The accuracy of the developed program was verified by comparing the calculated flying height of the coupled journal and thrust bearings in equilibrium (where the axial load generated by the FDBs was equal to the weight of a rotor) with the measured flying height at various rotating speeds (Jang et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure16-1.png",
        "caption": "Fig. 16. Wire drawing test.",
        "texts": [
            " However, when the implicit contact algorithm is preferred, smoothing allows reducing further the number of Newton\u2013Raphson iterations, so the CPU time is slightly decreased by smoothing: 1 h 41 min with smoothing and 1 h 44 min without. In conclusion, using a coarse tool discretization and combining the implicit scheme with the smoothing procedure provides the same accuracy than using a fine tool discretization with the standard approach but the computational time is strongly reduced: 1 h 41 min and 2 h 17 min respectively. The tool geometry and the wire finite element discretization are shown in Fig. 16 for the drawing process. The wire is pulled to pass through the die that reduces its diameter while increasing its mechanical properties. The material is considered as elastoviscoplastic. The contact between wire and die is assumed to be frictionless, which provides a good first approximation of the low friction of the actual process. The mesh size is approximately h = 0.75 mm and the tool discretization is about h = 0.5 mm. A detailed comparison between smoothed and unsmoothed simulations is presented in [24], where normals are computed by the consistency approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure5-1.png",
        "caption": "Fig. 5. The CHTC with rotational speed at 3L/min inlet flow rate.",
        "texts": [
            " As the rotational speed increases, the centrifugal and Coriolis effects make the turbulent flow separated near the shaft wall more intense and disorder, resulting in a higher CHTC. The oil of the recirculating hollow-shaft structure is introduced through the stationary tube and hits the bottom tail baffle to generate strong turbulence, so it has obvious advantages in the CHTC, especially in the case of high rotational speed. The CHTC of the recirculating hollow-shaft is 32% larger than that of the direct-through hollow-shaft when rotational speed is 5000rpm. As shown in Fig. 5, the CHTC at the bottom of the recirculating hollow-shaft gradually decreases with the increase of the rotational speed, and then gradually increases after the rotational speed reaches 3000 rpm. (a) 8.51e+02 6.81e+02 5.10e+02 3.40e+02 1.70e+02 0.00e+00 n=3000rpm (b) 3513 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. (a) n=2000rpm, (b) n=3000rpm, (c) n=5000rpm. As can be seen from Fig. 6, the recirculating flow I and recirculating flow II are generated when the oil is deflected by bottom tail baffle and top tail baffle",
            " In Section II, the phenomenon of a small CHTC at the bottom of the recirculating hollow shaft is analyzed when the rotational speed is 3000 rpm. In order to solve this problem, the bottom tail baffle shown in Fig. 3 is changed into a coneshape. Fig. 10 shows the velocity contour and vector diagram of the recirculating hollow-shaft with cone shape. Comparing Fig. 10 and Fig. 6(b), it can be seen that the cone angle facilitates the formation of a stronger toroidal vortex flow in the bottom area. As shown in Fig. 5(b), the CHTC at the bottom of the original structure is around 250 W/m2/K. However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization. Although changing the bottom to a cone-shape can improve the CHTC of the bottom area, the overall CHTC does not increase much. A simple optimization method is to gradually reduce the inner diameter of the bottom stationary cooling tube, so that it enhances the disorder of the bottom fluid, thereby improving the value of the overall average CHTC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001508_0954406214553983-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001508_0954406214553983-Figure10-1.png",
        "caption": "Figure 10. Calculation of U1\u00bd .",
        "texts": [
            "comDownloaded from The rotational matrix of bar 1 has the expression A1\u00bd \u00bc l4 cos sin ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p l3 sin 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p cos l3 sin ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p l4 cos ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p 0 l4 cos 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p l3 cos sin ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l2 3 sin2 \u00fel2 4 cos2 p sin 2 6666664 3 7777775 It results a1 b1 c1 2 64 3 75 \u00bc 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 \u00fe l24 cos 2 q A\u00bd 1 l3 l4 cos 0 2 64 3 75 \u00bc 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 \u00fe l24 cos 2 q 0ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 sin 2 \u00fe l24 cos 2 q l3 cos 2 64 3 75 , C1A n \u00bc i1 j1 k1 l6 0 0 a1 b1 c1 d1 e1 f1 2 64 3 75 \u00bc 0 l6c1 l6b1 2 64 3 75 N1f g \u00bc a1 b1 c1 d1 e1 f1 T a2 b2 c2 2 64 3 75 \u00bc 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 \u00fe l24 cos 2 q l3 l4 cos 0 2 64 3 75 C2A n \u00bc 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 \u00fe l24 cos 2 q i2 j2 k2 0 0 l7 l3 l4 cos 0 d2 e2 f2 2 64 3 75 \u00bc 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l23 \u00fe l24 cos 2 q l4l7 cos l3l7 0 2 64 3 75 N2f g \u00bc a2 b2 c2 d2 e2 f2 T : ii. Calculation of the matrix of restrictions U1\u00bd . From Figure 10 we have: \u2013 on the axis C1x we get the column 1 0 0 0 0 0 T because C1O1 i1 \u00bc 0; \u2013 on the axis C1y we obtain the column 0 1 0 0 0 l1 T because C1O1 j1 \u00bc l1k1; \u2013 on the axis C1z it results the column 0 0 1 0 l1 0 T because C1O1 k1 \u00bc l1j1. We deduce U1\u00bd \u00bc 1 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 l1 0 1 0 l1 0 0 0 2 6666666664 3 7777777775 iii. Calculation of the matrix of restrictions U2\u00bd . We use Figure 11 and find \u2013 on the axis C2x2 we obtain the column 1 0 0 0 l2 0 T because C2O2 i2 \u00bc l2j2; \u2013 on the axis C2y2 we get the column 0 0 0 l2 0 0 T because C2O2 j2 \u00bc l2i2; \u2013 on the axis C2z2 it results the column 0 0 1 0 0 0 T because C2O2 k2 \u00bc 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000822_j.proeng.2014.12.047-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000822_j.proeng.2014.12.047-Figure1-1.png",
        "caption": "Fig. 1. (a) 3-strut Simplex; (b) Top view of left and right modules.",
        "texts": [
            " Peer-review under responsibility of organizing committee of the XXIII R-S-P seminar, Theoretical Foundation of Civil Engineering (23RSP) Tensegrities consist of struts (marked on the drawings with thicker lines), which can be compressed, and cables which are solely tensioned. There are lots of regular and irregular tensegrity structures, including a series of typical modules, which can be used as a basis to build more complex systems. The following study focuses on a 3-strut Simplex module presented in Fig. 1a. Two types of modules were used: left and right (Fig. 1b). The following assumptions were used for calculation: strut lengths: 1.33 m, oblique cables lengths: 1.05 m, cables of smaller triangle: 0.58 m, cables of bigger triangle: 1.00 m. Struts were made of steel tubes with parameters: Es=210 GPa, As = 7.26 cm2, and cables of steel ropes with parameters: Es = 210 GPa, As = 2.01 cm2. The analysed module has one infinitesimal mechanism and a corresponding self-stress state, which stiffens the module. Cable removal causes the module to collapse. The further study considers a sample plate structure constructed from reversed 3-strut Simplex modules (described above), located in such a way that the left module is always joined with the right and vice versa (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.50-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.50-1.png",
        "caption": "Figure 4.50 Two inertias coupled with shaft",
        "texts": [
            " In particular, when choosing a motor, one assumes that the shaft size of the motor has been designed to accommodate the specified torque rating of the motor with a large safety margin. However, motor catalogs do not the specify the compliance rating of the shafts, as compared to the ratings of couplings which always specify both the torque and compliance rating (see Section 3.7). By ignoring compliance effects, a design can proceed well into the prototype phase, at which time oscillations are experienced and require a solution. Figure 4.50 shows a basic schematic of two inertias coupled together by a shaft: Jm represents the inertia of the motor rotor Jl represents the inertia of the load K1 is the stiffness of the shaft Bl is the viscous damping constant of the load This model will be explored in the following sections, although more complicated assemblies consisting of three and four inertias, such as an encoder or tachometer connected to the rear of the motor with the load connected to the front could be simulated. System Design 211 The equations expressing the torque relations for this model are: Tm = Jm\u03b8 \u2032\u2032 m + Bm\u03b8 \u2032 m + Jl\u03b8 \u2032\u2032 l (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002979_tii.2020.2986805-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002979_tii.2020.2986805-Figure8-1.png",
        "caption": "Fig. 8. The assembly platform and the objects. a) The CAD model where 1-3 are microscopes and 4-9 are six robot arms. b) The assembly platform, a close view of the assembly, and the objects.",
        "texts": [
            " This paper deals with inclined insertion with randomized postures, proposes withdrawal strategies to handle exceptional situations, and presents efficient insertion strategies considering the uncertainty introduced by adding a spring, of which the paper in [21] is incapable. Authorized licensed use limited to: University of Liverpool. Downloaded on July 27,2020 at 14:16:12 UTC from IEEE Xplore. Restrictions apply. 1551-3203 (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. We test the proposed method on the precision assembly platform, as shown in Fig. 8, which has three microscopes for visual feedback, two force sensors to acquire contact states, and six robot arms for multiple objects assembly. In this experiment, only two arms, two microscopes, and a force sensor are used for verification. The upper arm is equipped with three Suguar KWG06030-Gs, whose translational resolution is \u00b10.5\u00b5m, and a two-DOF manual tilt adjustment Sigma KKD25C. The lower arm has a Micos ES-100 with movement errors within 0.1\u00b5m, and KGW06050-L, KGW06075-L, and SGSP40yaw for rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003574_2374068x.2020.1835007-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003574_2374068x.2020.1835007-Figure3-1.png",
        "caption": "Figure 3. FDM patterns joined with sprue.",
        "texts": [
            " In order to build FDM patterns, 3-D solid models of the parts were fabricated employing SolidWorks 2010 software and exported as STL file to the FDM machine. Stratasys uPrint SE 3D printer was used for fabrication work. With layer thickness 0.254 mm, model temperature 25\u00b0C room temperature has been fixed for all the parts. According to designed experiments shown in Table 2, 18 patterns were built using ABS material. Nine sprues were also made with the help of FDM machine, and a liquid adhesive was used to join two patterns on single sprue as shown in Figure 3. ABS patterns produced by FDM process were finished by vapour smoothing process. The smoothing system used in this study was the FORTUS Smoothing station, Stratasys shown in Figure 4. It employs Decafluoropentane (30%) and Trans-di-chloro-ethylene (70%) smoothing fluid, having boiling point 43\u00b0C, as a solvent. This system is divided into two chambers, i.e. drying chamber and Smoothing chamber (330 \u00d7 406 \u00d7 508 mm). A confined smoothing chamber contains solvent and it is heated to 65\u00b0C by the heater at the bottom to create mixed vapour-liquid environment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002819_j.measurement.2020.107593-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002819_j.measurement.2020.107593-Figure5-1.png",
        "caption": "Fig. 5. Five types of sun gear faults.",
        "texts": [
            " The fixed-axis transmission box is supported by rolling bearings and has two-stage straight-tooth transmission. Vibration signals are acquired through the acceleration sensor and data acquisition instrument, and motor speed and brake are controlled through PC. Because the sun gear has high speed and meshes with other gears at the same time, it is easier to get faults. Therefore, the 5 faults of the secondary sun gear were simulated, including: broken gear, normal gear, gear with tooth root crack, gear with one missing tooth and wear gear as shown in Fig. 5. In the experiment, when output frequency of the motor is 25 Hz, 30 Hz, 35 Hz, 40 Hz and 45 Hz, fault simulation experiments are carried out for the 5 gear fault types, so 25 fault states in total are obtained as shown in Table 1. During the experiment, the brake load is set to 20 Nm, the sampling frequency is 12800 Hz, and each sample has 972,800 data point. In order to verify the validity of the fault diagnosis method for planetary gear under multi-operating conditions based on AEBoW model, the vibration signals obtained from the fault simulation experiment are put into the fault diagnosis model shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002940_s11044-020-09734-0-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002940_s11044-020-09734-0-Figure1-1.png",
        "caption": "Fig. 1 Four DoF golfer model relative to the inertial coordinate system at the initial condition",
        "texts": [
            " A global inertial frame was defined as a basis for the model\u2019s motion, with representation for the downrange target direction (X), the vertical (Y), and the perpendicular to the golfers stance (Z, following the right-hand-rule convention). The four DoFs of the golfer were provided by four revolute joints to allow for torso rotation, transverse flexion and adduction of the arm, supination and pronation of the forearm (alternatively, could represent the internal-external rotation of the shoulder), and ulnar and radial deviation of the wrist. This model is represented in Fig. 1. The human biomechanical model included four rigid body segments representative of the golfer\u2019s single upper arm, forearm, hand, and torso. To provide a realistic golfer stance, the torso segment was constrained to rotate about an axis 35\u25e6 to the vertical, in which a zero degree rotation was defined as the approximate address position of the golfer. The swing plane of the arm was rotated 70\u25e6 from the horizontal [7], which allowed the arm to sweep downwards towards the ball during impact. A zero degree angle of the shoulder approximated the address position of the golfer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002406_j.ifacol.2016.09.007-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002406_j.ifacol.2016.09.007-Figure1-1.png",
        "caption": "Fig. 1 Terminal homing engagement geometry",
        "texts": [
            " Inspired by the previous work, considering the FNTSM which can combine the relationship between the error and error rate of LOS and easing the coupling relationship among the chattering amplitude, disturbance compensation and convergence rate of sliding variable, a new fast nonsingular terminal sliding mode to attenuate the chattering guidance (SMACG) law is proposed. 2. MODES AND DISTURBANCE OBSERVER 2.1 Mode Description The two dimensions planar terminal homing engagement geometry between a guided missile and the manoeuvring target is depicted in Fig. 1 , where the subscripts M and T denote the guided missile and the manoeuvring target,  and R the LOS angle and the guided missile- manoeuvring target relative range, MV and TV the guided missile and manoeuvring target velocity, M and T the guided missile and manoeuvring target flight path angle, MA and TA the guided missile and maneuvring target acceleration, which are normal to their corresponding velocities, respectively. Assumption 1: the guided missile and manoeuvring target are point masses moving in the 2D plane with constant velocities, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000062_raee.2019.8886946-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000062_raee.2019.8886946-Figure2-1.png",
        "caption": "Fig. 2. Conventional 12S-10P E-Core SFPMM (a) 2-D Cross sectional view (b) Flux distribution",
        "texts": [
            " Moreover, PM usage is same which means that due to extra mechanical accessories, mechanical adjustor topologies results in increase of cost. SFPMM exhibits outstanding electromagnetic performance, however increase in PM usage further increase machine cost. To overcome PM usage in SFPMM, an overview is carried out in [6] showing different topology of SFPMM with reduced PM usage whereas E-core SFPMM is introduced in [11] which reduced PM usage to half and term as conventional 12/10 stator slot/rotor pole (12S-10P E-Core SFPMM) as shown in Fig. 2(a). However, flux leakage is neglected and not taken in consideration. Flux distribution of 12S-10P E-Core SFPMM is shown in Fig. 2(b). Analysis reveals that despite of reduce in PM usage, still there are significant flux leakage. A similar performance evaluation of conventional 6S-10P SFPMM with E-Core and C-Core is carried out in [12] that reduce PM usage and enhance slot area. Moreover, its performance is compared with conventional 12S-10P E-Core SFPMM. Fig. 3(a) and Fig. 3(b) shows 2-D cross sectional view and flux distribution of 6S-10P E-Core SFPMM whereas Fig. 4(a) and Fig. 4(b) shows 2-D cross sectional view and flux distribution of 6S-10P C-Core SFPMM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000253_iros40897.2019.8967789-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000253_iros40897.2019.8967789-Figure6-1.png",
        "caption": "FIGURE 6. Thread-per-Edge LDPC decoder thread-parallelism approach.",
        "texts": [
            " allowed decoding throughputs of dozens to hundreds ofMbit/s [23]\u2013[25], combining a graphics language approach with Caravela streams [97], as illustrated in Fig. 5. thread-per-edge (TpE) corresponds to the strategy with the finest granularity, which brings upon the LDPC decoder designer a granularity tradeoff. For the one, if consecutive threads deal with the update ofmessages belonging to consecutive nodes in the Tanner graph, then there is a high exposure of both spatial and temporal data locality, as Fig. 6 illustrates. For the other, most GPU-based LDPC decoders that exploit this granularity have been developed for pre-Fermi or Fermi architectures that do not have a caching mechanism available to threads for computation [6]\u2014it exists only for off-chip VOLUME 4, 2016 6709 memory transaction. For instance, locality is automatically explored by the x86 cache system in heavily SIMD-based LDPC decoders [75], leading to over 90% cache hit rates that maximize the LDPC decoder bandwidth. Alas, this is not the case in the many-core-based decoders utilizing this strategy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure21-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure21-1.png",
        "caption": "Figure 21. Speed field contours for 600 rpm.",
        "texts": [],
        "surrounding_texts": [
            "Table 2 and the graph (Figure 19) show the calculations of the maximum speed of the velocity fields for a fixed drum diameter of 650 mm and different values of the drum rotation speeds. The diameter and rotation speed of the beater are equal to the initial values. [22] ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0000577_transducers.2015.7181238-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000577_transducers.2015.7181238-Figure6-1.png",
        "caption": "Fig. 6: The completely fabricated FO-SPR sensor",
        "texts": [
            " To deposit a uniform and homogeneous metal layer on the circumferential surface of the fiber optic, a novel metal deposition method combining the turntable revolving and the fiber optic rotating was brought forward. The fiber optic was mounted to a rotator, as shown in Fig. 5(b). With the clamped rotator rotating and the turntable revolving simultaneously, the metal particles evaporated from the bottom of the deposition system could be deposited uniformly to the cylinder of the fiber optic. After the metal deposition, an SMA905 connector was installed on both sides of the fiber optic to couple the light. Fig. 6 shows the FO-SPR sensor fabricated. The FO-SPR sensor was immersed in a solution (1:1:5 mixture of ammonia, H2O2 and H2O) at 75oC for 15 minutes, followed by rinsing with plenty of ultrapure water and drying with a nitrogen stream. And after that, it was prepared for polymer assembly. In this paper the borate polymer was synthesized, and then immobilized onto the gold surface by the layer-by-layer (LBL) self-assembly technique. This technique is an alternative deposition method using weak interactions (electrostatic attraction between ions, hydrogen bond, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000189_j.matpr.2019.11.251-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000189_j.matpr.2019.11.251-Figure10-1.png",
        "caption": "Fig. 10. Mesh pattern of the designed arm in ABAQUS to perform bending simulations.",
        "texts": [
            " C3D8R is a three-dimensional brick element with reduced integration, capable of improving computational efficiency and have only translational degree of freedom at the nodes. These elements are capable to capture both mechanical stresses and strains. The present investigation has not accounted cohesive layer modelling but considered the effect of same in carbon/glass composite laminate. Further, to reduce modelling complexities and computational time, only half of the composite was modelled, illustrated in Fig. 10. The integration between the balsa wood (core) and carbon/glass composite (facesheet) was achieved by TIE constraint to illustrate as resin epoxy. It was assumed that the TIE link between consecutive plies is rigid. The size of an element considered for mechanical analysis of the composite was 5 mm 5 mm 5 mm based on the mesh convergence study carried out. The results from finite element simulations were compared for flexural strength and energy absorbed before the failure. The stress-strain curve of sandwich composite cantilever in bending action highlights that both balsa carbon composite and balsa glass composite satisfy the failure limit of 318 MPa for the MRS"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000934_1.4024781-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000934_1.4024781-Figure1-1.png",
        "caption": "Fig. 1 Outline of an assembly of a gasket and two flanges",
        "texts": [
            " Compared to the thin polymer film method that merely detects the contact marks, the present method showed clear images of the entire gasket surfaces that helped gain an understanding of the real contact situation between the flanges and gasket. From the observation results, we could easily determine the critical contact pressure where the leakage paths in the radial direction disappear. Over the critical contact pressure, the leakage rates obtained from the experiments were in good agreement with the calculated values under an assumption of laminar flow. We reported a method for prediction of leakage rates based on the visualization of real contact area by the polymer film method [13]. Here, these results are briefly explained. Figure 1 shows an outline of an assembly of a gasket sandwiched by two flanges. The polymer film method was used to measure the real contact area between two solid surfaces by insertion of a thin polymer film between them onto the surface of which contact marks were transferred. In this case, the polymer film was inserted between the flange and the gasket. As the contact marks on the polymer film appear dark under an optical microscope, the real contact area can be measured through image processing, such as thresholding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000999_j.proeng.2014.12.629-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000999_j.proeng.2014.12.629-Figure1-1.png",
        "caption": "Fig. 1 Novlit-3 tail-sitter VTOL MAV",
        "texts": [
            " Peer-review under responsibility of Chinese Society of Aeronautics and Astronautics (CSAA). Keywords: adaptive control; Dynamic inversion; Tail-sitter MAV; Hover The capability of vertical take-off and landing (VTOL) enables Micro Aerial Vehicle (MAV) to execute lowspeed search in limited space or even stare surveillance. Tail-sitter configuration tends to be an appropriate solution for VTOL MAV since it achieves transition flight between high efficient cruise and prop-hang hover without involving additional tilting mechanism. Novlit-3 (Figure 1) is a tail-sitter VTOL MAV with 60cm wing span and 700g take-off weight, which has a novel X shaped wing layout with a single propeller. Its non-orthogonal wing layout and the corresponding control surfaces arrangement provide extra lateral-directional stability in cruise and * Corresponding author. Tel.: +86-139-9126-0177; fax: +86-29-88493672. E-mail address: wangjin_nwpu@163.com 015 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (http://creativecommons"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure17-1.png",
        "caption": "Fig. 17: Tube-to-tube welding simulation",
        "texts": [
            " 12 and Fig. 13 respectively. All simulation results can show the effectiveness of our proposed method. The model parameters of tube-to-plate and welding discrete seam are shown in Fig. 14. Simulation results are drawn as follows. Joint angels of the master and slave robot are shown in Fig. 15 and Fig. 16 respectively. All simulation results can show the effectiveness of our proposed method. In this simulation, the model parameters of tube-to-tube and the workpiece coordinate system are shown in Fig. 17(a). And the workpiece coordinate system, the master robot tool coordinate system and the master robot flange coordinate system are coincident with each other. In addition, the workpiece coordinate system is fixed to and moves along tube-to-tube plate. The welding seam discrete points of tube-to-tube welding are shown in Figure 17(b). In this paper it is discretized into 248 welding points. The blue line represents the normal vector (axis Z) of each welding point. In the simulation space, the search area is shown in Fig. 18. We set the search length as 440 mm, 450 mm and 453.33 mm along the axis X, Y and Z. So we can get 60 experimental points. In the evaluation function (9) of each experimental point, we set wj as 1 and the sampling interval as 0.1s. According to the length of the weld (617.51 mm) and the welding speed (10 mm / s), the number of sample points comes to 618"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000657_j.ijmecsci.2015.10.007-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000657_j.ijmecsci.2015.10.007-Figure5-1.png",
        "caption": "Fig. 5. Analytical model of the rectangular thin film secondary wrinkling under shear.",
        "texts": [
            " Meanwhile, the dimensionless factor \u03bbcr lIIcr corresponds to the value of the critical load of thin film secondary wrinkling. The secondary wrinkling will occur when the compressive stress equals to \u03c3II cr , which is expressed as follows. \u03c3II cr \u00bc Eh2 12 1 v2 b2 \u03c0 lIIcr !2 m2\u00fen2lII2cr 2 m2 cn2lII2cr \u00f017\u00de Based on the bifurcation theory with the consideration of rectangular thin film shear wrinkling, a critical load prediction method of the thin film secondary wrinkling can be proposed. The analysis model of the rectangular thin film secondary wrinkling under shear is shown in Fig. 5. In Fig. 5, L1 and L2 are respectively the length and width of the rectangular thin film model, while \u03b1 and \u03b2 represent the wrinkling direction and the texture direction, respectively. Moreover, \u03bb1 and \u03bb2 are the wrinkling half-wavelength in \u03b1-direction and \u03b2-direction, respectively. \u03b8 and \u03b3xy are respectively the wrinkling angle and the shear strain. Because the displacement u1 is far less than L2, the shear strain can be expressed as \u03b3xy \u00bc u1=L2. Meanwhile, the relationship between the wrinkling half-wavelength \u03bb1 and the wrinkling angle \u03b8 can be expressed as \u03bb1 \u00bc L2= sin \u03b8. According to the conditions of the boundary and load (Fig. 5), the maximum principal stress \u03c3\u03b1 and the minimum principal stress \u03c3\u03b2 can be obtained as follows. \u03c3\u03b1 \u00bc E 2 1 v\u00f0 \u00de v1 L2 \u00fe E 2 1\u00fev\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 L2 2 \u00fe u1 L2 2r \u03c3\u03b2 \u00bc E 2 1 v\u00f0 \u00de v1 L2 E 2 1\u00fev\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi v1 L2 2 \u00fe u1 L2 2r ; \u03b3xyo\u03b3Ixy cr 8>>< >>: \u00f018\u00de Herein, \u03b3Ixy cr is the critical shear strain of the thin film primary wrinkling. In addition, the wrinkling angle \u03b8 can be expressed as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000506_978-3-7091-1379-0_2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000506_978-3-7091-1379-0_2-Figure3-1.png",
        "caption": "Figure 3: Type 1 singularities",
        "texts": [
            " Matrices (4) can be used for the analysis of all three types of singularities (Gosselin and Angeles, 1990). It was found that the Type 1 singularity (occurs when matrix x = 0;x = \u22122l2; y = 0; y = \u22122l2; z = 0; z = \u22122l2. (5) y = \u2212z tan \u03b81; z = \u2212x tan \u03b82;x = \u2212y tan \u03b83. (6) It was also found that Type 3 singularity (when both A and B are singular) occurs when at least one of conditions (5) is satisfied. One can notice that conditions (5) correspond to the limit points of the [2l2; 0] interval. Conditions (6) correspond to the situation when all links of any leg lie on the same plane (Figure 3a). Thus, it is easy to conclude that the conditions (6) can be satisfied only at the edge of the workspace. In Figure 3b the singular surfaces for (6) are shown. For Type 2 singularities (occurs if matrix A is singular) no analytical conditions were found using Jacobian analysis. Numerical analysis of the workspace proves that there is no point within the workspace where det(A) is zero. Moreover, at every analyzed point det(A) is less than zero. We have also carried out the analysis based on the screw theory and obtained the same analytical conditions for Type 1 and Type 3 singularities, as when we used previous method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.7-1.png",
        "caption": "Figure 3.7 Wound field versus permanent magnet motor",
        "texts": [
            " System Components 39 Beginning in 1930, a powerful permanent magnet material, Alnico (aluminum, nickel, cobalt) was created that has, to a large extent, replaced the wound stator poles and resulted in the development of today\u2019s permanent magnet brush motor. In the 1960s a relatively inexpensive permanent magnet material, Hard Ferrite, a compound of various ceramics and iron oxide was developed, helping to create a class of low cost brush motors. A cross-section of a typical four pole permanent magnet brush motor is shown in Figure 3.6. This is illustrated in Figure 3.7. Regardless of whether the field is created by current-carrying coils mounted on salient poles or by permanent magnets, armature control assumes that the field is constant and all control is dependent on control of the armature current. 40 Electromechanical Motion Systems: Design and Simulation The relation between the applied current and the resulting torque is expressed as: T = Kt Ia (3.3) where Kt is the torque constant in N m A\u22121, Ia is the applied current in A and T is the resulting torque in N m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003563_j.ijleo.2020.165776-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003563_j.ijleo.2020.165776-Figure9-1.png",
        "caption": "Fig. 9. Distribution of residual stress field of the valve seat.",
        "texts": [
            " However, due to the shortening of the time for laser irradiation of the material, the uneven temperature distribution led to an increase in the temperature gradient in the phase transformation area near the molten pool. Therefore, there was a competitive relationship between the influence of temperature and temperature gradient on stress. Temperature distribution, temperature gradient distribution and structural constraints affect the magnitude and distribution of stress. When the laser scanning speed is 3 mm/s, the residual stress field distribution of laser remelting of valve seat is shown in Fig. 9, and the residual stress distribution of path 3 and path 4 is shown in Fig. 10. During the laser remelting process, the thermal stress generated by the high temperature exceeded the yield limit of the material, the material of the molten pool was plastically deformed, and the residual stress was generated during the cooling process. The residual stress in the cone area of the valve seat scanned by laser was tensile stress. The larger residual stress was mainly concentrated in the remelting zone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001898_s00542-015-2527-2-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001898_s00542-015-2527-2-Figure2-1.png",
        "caption": "Fig. 2 The CAD model of HDD",
        "texts": [
            " The slider dynamic simulation could consider the nonlinear air bearing stiffness from the flying height of the slider and analyze the slider attitude. The slider dynamic analysis result was used to investigate the anti-shock performance. setup Accelerometer Measuring point Drop tester Jig LDV Fig. 7 The comparison between experiment and simulation results 0 0.5 1 1.5 2 2.5 3 -0.4 -0.2 0 0.2 Time, msec D is pl ac em en t, m m Experiment Simulation 1st peak 2nd peak 1 3 1 3 A finite element method (FEM) model of a 2.5-inch HDD was constructed to predict the slider and disk dynamics with ramp\u2013disk contact. Figure 2 shows the components of the HDD. The FEM model included the base, spindle motor, disk, HGA, slider, cover, ramp, e-block, slider and ABS spring. A transient shock was simulated with the FEM model to extract the vertical force and moment between the slider and disk. The slider and disk are connected by air bearing springs. A real air bearing system can be described as a nonlinear spring in the operational shock situation. However, the air bearing spring can be simplified as a linear spring to speed up the simulation of the structural dynamic response (Liu et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure4-1.png",
        "caption": "Fig. 4 The mesh map of the part of the gear pump gear teeth in cycle one",
        "texts": [
            " According to the data in Table 1, the following expression can be expressed as follows: Dx x6 1 \u00bc 2H~6 11D~11 1 \u00f032\u00de The damage coefficient vector [d1, d2, d3, , d11]T can be obtained through solving the expression (32), which is listed as follows: D~ \u00bc \u00bdd1; d2; d3; d4; d5; d6; d7; d8; d9; d10; d11 T \u00bc \u00bd0:0005; 0:0027; 0:0024; 0:0017; 0; 0; 0; 0; 0:0023; 0:0002; 0:0003 T Where d3 and d9 are \\0, and they are set as 0, and the expression (32) is solved again, and the new damage coefficients are obtained as follows: D~ \u00bc \u00bdd1; d2; d4; d5; d10; d11 T \u00bc \u00bd 0:00036; 0:0027; 0:0014; 0:0032; 0:0018; 0:00021 T Where d1, d2, d5 and d10 are\\0, and they are set as 0, and the following results are as follows: D~ \u00bc \u00bdd4; d10 T \u00bc \u00bd0:0012; 0:0003 T According to the final results, the elements of the damage matrix are all positive value, therefore, the zonesP e4 and P e10 can be predicted as cracked elements. The cracked zones P e4 and P e10 are extracted, and they are divided into three zones, respectively, and element division in two zones is carried out again, and the wavelet finite element meshes are shown in Fig. 4. The element effect matrix of H~P e4 and H~P e10 is obtained through interpolating in the solving domain, and the new response matrix H~1 can be obtained through combining H~P e4 with H~P e10 , and the damage coefficient matrix can be expressed as follows: D~ \u00bc \u00bdd11; d12; d13; d14; d15; d16 T \u00bc \u00bd 0:018; 0:3126; 0:0195; 0:0538; 0:0672; 0:0336 T The component which is negative value can be set as 0, and the new damage coefficient matrix can be expressed as follows: D~1 \u00bc \u00bdd12; d15 T \u00bc \u00bd0:075; 0:023 T The precision error is set as n = 10, and the same pro- cedure is carried out again"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002852_1045389x20906474-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002852_1045389x20906474-Figure1-1.png",
        "caption": "Figure 1. Magnetic screw pump concept with flights made of magnetized MRF (left) compared to a typical screw pump with metal flights (right).",
        "texts": [
            " The first presents the flow tests results and confirms the pumping action of the magnetic screw pump. The second subsection introduces a semi-empirical yield screw pump model to represent the experimental flow tests results. The third subsection presents the durability tests results, where the effect of circulation on MRF durability is investigated. Finally, the conclusion summarizes the main results of the paper and discusses future work. The magnetic screw pump model proposed in the article is inspired by a typical screw pump as shown on the right of Figure 1 (e.g. for plastic extrusion), but instead of having solid flights, the flights are generated by 3D structures of MRF formed under the effect of a concentrated magnetic field (left of Figure 1). Figure 2 shows a drum-type MR clutch with a grooved drum (dark green) and a magnetic circuit (blue). Helical or spiral grooves are machined into the drum and act as a screw thread. Grooves create a local reluctance in the magnetic field path and cause the flux lines to go around the grooves as shown on the magnetic field distribution (Figure 2). The MRF contained in the grooves is unmagnetized (or slightly magnetized), while the portion of MRF located between the grooves (on the flat surfaces) remains strongly magnetized and maintains the clutch\u2019s primary function of transmitting torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000750_j.jweia.2014.11.004-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000750_j.jweia.2014.11.004-Figure4-1.png",
        "caption": "Fig. 4. Ricochet of a debris.",
        "texts": [
            " Thus, it is necessary to include the ricochet effects when predicting the debris travel distance. In the DeThrow II programme, the ricochet criterion suggested by Xu et al. (2014) which was thoroughly verified experimentally and numerically was adopted. Toward this end, the following relationships for the debris ricochet are adopted Xu et al. (2014). For \u03b8iZ311Vf \u00bc 0; \u03b8f \u00bc 0 \u00f015\u00de Otherwise \u03b8f \u00bc 0:15 \u03b8i\u00fe121; Vf \u00bc \u00f00:72 \u03b8i \u03c0=180 0:0016\u00deVi \u00f016\u00de In Eqs. (15) and (16), Vi and \u03b8i are the impact velocity and the impact angle of the debris, respectively (Fig. 4) while Vf and \u03b8f are the post impact velocity and ricochet angle of the debris. In order to predict the debris flight trajectory, the equations are numerically solved by the fourth-order Runge\u2013Kutta algorithm (Sewell, 2005) and the general flow chart used by the DeThrow II programme is shown in Fig. 5. The tolerance and the minimum time step are set to be 0.001 m and 0.001 s, respectively. Based on a series of numerical simulations of the initial breakup of above ground magazine, Fan et al. (2014) plotted each occurring debris according to its length ratio as a probability density function (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001312_j.jfranklin.2013.09.017-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001312_j.jfranklin.2013.09.017-Figure2-1.png",
        "caption": "Fig. 2. Collision geometry.",
        "texts": [
            " The points \u00f0xp; yp\u00de and \u00f0xe; ye\u00de are the current coordinates; Vp and Ve are the velocities; ap and ae are the lateral accelerations of the pursuer and the evader respectively; \u03c6p and \u03c6e are the respective angles between the velocity vectors and the reference line of sight; and y\u00bc ye yp is the relative separation normal to the initial line of sight. If the velocities Vp and Ve are constant, then the aspect angles \u03c6p and \u03c6e can be expressed as follows: \u03c6p\u00f0t\u00de \u00bc \u03c6p\u00f0col\u00de \u00fe \u0394\u03c6p\u00f0t\u00de; \u03c6e\u00f0t\u00de \u00bc \u03c6e0 \u00fe \u0394\u03c6e\u00f0t\u00de; \u00f01\u00de where the \u201ccollision course values\u201d \u03c6p\u00f0col\u00de and \u03c6e0 are defined by the following equation: Vp sin \u03c6p\u00f0col\u00de \u00bc Ve sin \u03c6e0 \u00f02\u00de (see Fig. 2). If both aspect angles \u03c6p and \u03c6e remain close to the \u201ccollision course values\u201d, i.e. \u0394\u03c6p\u00f0t\u00de as well as \u0394\u03c6e\u00f0t\u00de remain small, then the original nonlinear engagement model can be linearized. In such a case the final time tf can be prescribed and calculated as tf r0=\u00f0Vp cos \u03c6p\u00f0col\u00de\u00fe Ve cos \u03c6e0\u00de, where r0 is the initial distance between the objects. Moreover the original 3D geometry can be decoupled into two identical planar ones in perpendicular planes [21]. The planar linearized model, including the first-order dynamics with time constants \u03c4p and \u03c4e and assuming constant maximal lateral acceleration commands amax p and amax e (respective maneuverabilities), is described [22] by the differential equation _x \u00bc Ax\u00fe bu\u00fe cv; \u00f03\u00de with the state vector x\u00bc \u00f0x1; x2; x3; x4\u00deT \u00bc \u00f0y; _y; ae; ap\u00de, and A 0 1 0 0 0 0 1 1 0 0 1=\u03c4e 0 0 0 0 1=\u03c4p 2 66664 3 77775; \u00f04\u00de bT \u00bc \u00f00; 0; 0; amax p =\u03c4p\u00de; \u00f05\u00de cT \u00bc \u00f00; 0; amax e =\u03c4e; 0\u00de: \u00f06\u00de The initial state is x0 \u00bc \u00f00; x20; 0; 0\u00deT , where x20 \u00bc Ve\u03c6e\u00f00\u00de Vp\u03c6p\u00f00\u00de"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001534_0954410012472292-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001534_0954410012472292-Figure2-1.png",
        "caption": "Figure 2. CATIA measure inertia function.",
        "texts": [
            " Then, the dynamic analysis of the damaged UAV by conducting numerical simulations with its longitudinal and lateral/directional models is discussed. Lastly, conclusions and future works are given. A flying-wing type single jet UAV was developed for modeling and analysis of the damaged airplane. The partial right-wing damage is assumed to have a loss of area moments of 22%, 60%, and 78.4%. Figure 1 shows the geometry of the flying-wing type UAV considered in this research, whose body/span lengths and sweep angle are 1.883m/1.98m and 43 , respectively. The partial wing loss causes the UAV to reduce its total mass. Figure 2 represents the trend of the mass decrease with respect to the percentage of wing damage, where the total mass reduces to 12% when the wing area moment is damaged by 78.4%. This is found from using CATIA measure inertia function, Table 1. Test condition of wind tunnel test. Damage magnitude 0% 22% 60% 78% Test range Angle of attack 5 to \u00fe25 5 to \u00fe25 5 to \u00fe25 5 to \u00fe25 Sideslip angle 30 to \u00fe30 30 to \u00fe30 30 to \u00fe30 30 to \u00fe30 Aileron deflection 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 Elevator deflection 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 Rudder deflection 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 20 to \u00fe20 at UNIV OF PITTSBURGH on March 11, 2015pig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure10-1.png",
        "caption": "Fig. 10. The velocity vector of the recirculating hollow-shaft with cone angle at an inlet flow rate of 3L/min and a rotational speed of 3000rpm.",
        "texts": [
            " It is noted that friction loss of recirculating hollow-shaft is higher than direct-through hollow-shaft. III. STRUCTURE OPTIMIZATION OF RECIRCULATING HOLLOW-SHAFT From the analysis of the CHTC above, it is noted that the CHTC in the middle region is small and uniform compared to bottom wall and top wall of the recirculating hollow-shaft. In Section II, the phenomenon of a small CHTC at the bottom of the recirculating hollow shaft is analyzed when the rotational speed is 3000 rpm. In order to solve this problem, the bottom tail baffle shown in Fig. 3 is changed into a coneshape. Fig. 10 shows the velocity contour and vector diagram of the recirculating hollow-shaft with cone shape. Comparing Fig. 10 and Fig. 6(b), it can be seen that the cone angle facilitates the formation of a stronger toroidal vortex flow in the bottom area. As shown in Fig. 5(b), the CHTC at the bottom of the original structure is around 250 W/m2/K. However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization. Although changing the bottom to a cone-shape can improve the CHTC of the bottom area, the overall CHTC does not increase much. A simple optimization method is to gradually reduce the inner diameter of the bottom stationary cooling tube, so that it enhances the disorder of the bottom fluid, thereby improving the value of the overall average CHTC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003367_s00170-020-05832-7-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003367_s00170-020-05832-7-Figure7-1.png",
        "caption": "Fig. 7 Step 4 temperature field distribution",
        "texts": [
            " During the simulation, Surf152 surface convection heat exchange unit should be introduced into the grinding area of tooth surface. According to Jaeger\u2019smoving heat source theory and the convection heat transfer model of grinding fluid on the tooth surface proposed by Des Ruisseaux, the moving heat source model of cycloid gear tooth surface is established, as shown in Fig. 5. Figure 6 shows the loading situation of grinding fluid convective heat transfer load. The temperature field distribution of the tooth surface in the fourth load step is shown in Fig. 7, when the grinding process parameter is selected as follows: the grinding wheel diameter 264 mm, the width of cycloid gear 20 mm, Vs = 40m/s, Vw = 3.5m/min, and Ap = 0.025mm. The simulation calculation under includes 8 load steps, as shown in Fig. 8. The highest temperature of the tooth surface in this load step is 764.748 \u00b0C, which locates near the tooth top. According the view of heat transfer, this phenomenon is caused by the protruding of the profile. There are relatively few materials available for heat conduction in this place"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure4-1.png",
        "caption": "Fig. 4. Novel homopolar-type MB unifying the four C-shaped cores. (a) Appearance of a novel homopolar-type MB. (b) Magnetized direction of permanent magnets.",
        "texts": [
            " In addition, it is very difficult to increase the stator inner surface accuracy. 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In order to realize the stable magnetic suspension and the low loss, it is important that the stator inner accuracy be highly maintained. This paper proposes a novel homopolar-type MB structure unifying the four C-shaped cores for high output, low rotor iron loss, and simple assembly. Fig. 4(a) shows the novel homopolar-type MB structure. The stator core of the novel homopolar-type is divided into stator bulk cores and stator laminated cores. The windings are wound to the stator bulk cores. The stator-laminated cores are made up of unifying teeth parts, in which the magnetic flux flows to the radial direction in the four separated C-shaped cores, as shown in Fig. 3. Therefore, assembling four separated C-shaped cores is not necessary. In addition, the width between the magnetic poles, shown in Fig. 4(b), easily becomes small by the unified stator laminated cores. It is noted that the leakage flux between the magnetic poles is blocked by inserting small magnets between the magnetic poles, as shown in Fig. 4(b). Since the novel homopolar-type MB has no large magnets for generating bias flux, the magnet volume compared with that in the general homopolar-type can be decreased. Thus, the cost of the novel homopolar-type MB can be reduced. Fig. 5 shows the procedure of a stator assembly for the novel homopolar-type MB in detail. The stator-laminated cores are inserted into a nonmagnetic material case from the axial direction, as shown in Fig. 5(a). In addition, small magnets are inserted to the holes of the stator-laminated cores",
            " The magnetic flux \u03c6x+ is generated by the current ix+ flowing in the Nx+ winding, as shown in Fig. 6. Then, the magnetic attractive force Fx+ toward the positive direction in the x-axis is generated in the gaps between the stator laminated cores and the rotor laminated cores. Similarly, Fy+ toward the positive direction in the y-axis is generated. Thus, the suspension force Fb+ can be generated toward the positive direction in the b-axis by a vector sum of Fx+ and Fy+. Based on the principle of generating the suspension force in Section III-B, the novel homopolar-type MB, shown in Fig. 4(a), is evaluated with the 3D-FEA. In addition, in order to compare with the general types, shown in Figs. 1 and 2, the heteropolar-type MB and the general homopolar-type MB are prepared. These types are predesigned with the 3D-FEA to generate the maximum suspension force in the same volume size, respectively. Table I shows the specifications of the heteropolar-type MB, the general homopolar-type MB, and the novel homopolar-type MB. From Table I, the magnet volume of the novel homopolar-type MB is decreased to only 7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002852_1045389x20906474-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002852_1045389x20906474-Figure3-1.png",
        "caption": "Figure 3. Picture and partial cross section of a grooved drum with two grooves and a pitch of 12.7 mm with grooves dimensions. The blue arrow represents the MRF path when the drum is installed in an MR clutch.",
        "texts": [
            " This solid MRF structure follows a helical pattern and transports the unmagnetized MRF as the clutch rotates. A finite element model run on the the FEMM (Finite Element Method Magnetics) software confirms that a low magnetic field in the 0 to 0.1 Tesla range is found inside the grooves while the remainder of the MRF gap is at a strong field of around 1 Tesla. The magnetic circuit of the model is made of 1018 steel and the coil is made of 220 turns of 25 AWG magnet wire. The current flowing through the coil is 4 A. Figure 3 presents an actual MR clutch drum used in this work. Helical grooves resembling a screw thread are machined on the inner and outer surfaces of the drum. Figure 3 also shows the partial cross section of the grooved drum with two 12.7-mm pitch grooves on each drum surface. The grooves are of triangular shape with a 60 angle and made on a lathe with a standard threading tool. This groove shape and dimensions is chosen because it provides an unmagnetized area into the groove and it can be manufactured easily. The proposed pump design requires no additional external parts, which is interesting from a reliability and complexity standpoint. A simple analytical model is derived from typical screw extruder theories (Birley et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001860_robio.2015.7419705-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001860_robio.2015.7419705-Figure4-1.png",
        "caption": "Fig. 4: Constant bearing strategy. P and T are the current positions of the pursuer and the target, respectively.",
        "texts": [
            " The origin (O) is the starting point for the constant bearing interception trajectory and it is the point at which the reference bearing is stored. Again, the velocities of the pursuer (Vp) and the target (Vt) are constant for a particular duel. The motion of the target is as in Section IIIB. However, the pursuer no longer simply moves directly towards the target\u2019s line of sight (LOS). Instead, the pursuer attempts to move along a linear trajectory in such a way as to hold the absolute bearing of the target constant. Consider the angle between the instantaneous LOS (PT) and the reference frame in Fig. 4, this angle, denoted by the bearing (\u03b8), is computed by (8). \u03b8e = arccos ( P\u0302 T \u00b7 y\u0302 ) (8) The instantaneous bearing (\u03b8) is compared with the initial bearing (\u03b8o), which is the reference angle that will be used to keep the absolute bearing constant. As the velocity is unchanging throughout each duel, the absolute bearing angle is held constant by varying the trajectory angle (where \u03b1 is the current trajectory angle of the pursuer and \u03b1o is the previous trajectory angle.), i.e. if \u03b8 > \u03b8o then \u03b1 increases and visa versa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000849_powereng.2013.6635624-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000849_powereng.2013.6635624-Figure2-1.png",
        "caption": "Fig. 2. Three-dimensional slot wedges with axial and tangential gaps",
        "texts": [],
        "surrounding_texts": [
            "The generator and the exciter have a very high mass of inertia. During the starting process it is possible that the rotor begins to run asynchronously, which causes a high thermal loading on sensitive components. By using the method of finite elements this problem is studied in different cases. Another part, which gets observed, is the general behaviour and the electrical and mechanical stresses of the system.\nI. INTRODUCTION\nThe federal government in Germany decides to shut down all nuclear power plants until the year 2022. Old plants are already closed. Since the expansion of renewable energies and the changing situation of electricity generation there needs to be voltage compensation near Frankfurt to secure the voltage stability of the grid in this industrial location. The Federal Network Agency placed the order to modify a generator to a synchronous condenser.\nNormally, very large synchronous generators are coupled to a driving shaft and start with steam turbines as main drives. These types of generators are separate excited synchronous machines and can be started with a converter as well, like every synchronous machine. The start-up can be only done with reduced voltage amplitudes, since there is no converter on the marked with an apparent power of 1.5 GVA. Even with reduced amplitudes there are some constructive details, which differentiated large synchronous generators from variable speed drives.\nEdward D. Goodman and Thomas H. Barton show in early studies [1], that there is an oscillating torque during an asynchronous run. They have a close look at the start-up by using the damper windings. The operating torque can be divided into different parts. The Fourier transformation shows a accelerating part, which caused by induced currents in the\ndamper bars, an oscillating part and parts of higher harmonics.\nThe field winding is affixed at the rotor by slot wedges. These wedges are used as damper windings, too. They were mainly needed to stop oscillations of the magnet wheel during load cycles or during similar situations. If the generator begins to run asynchronously, then there will be high currents induce in the slot wedges, which result in a thermal load. It is necessary to calculate the temperature in the wedges to avoid damage from the insulation of the field winding and other sensitive components.\nThe last aspect, that needs to be considered, is the amplitude of the phase winding current. During the operation, the magnet wheel turns over 180 degrees electrical. This leads to increased currents.\nTore Petersson and Kjell Frank looked at this topic in",
            "1972 [3]. They study about starting synchronous machines with frequency converters and make a statement about the thermal load of damper bars. Their analysis is based on the two reaction model by R.H. Park, which is based on certain assumptions and simplifications. With the help of up to date finite elements and time stepping analysis it is possible to get a closer look on this topic.\nThe electro-mechanical model is built up with the software ANSYS Mechanical APDL. It is a two dimensional crosssection of the generator. Figure 1 shows the whole FE-Model. By using symmetrical properties the model can be reduced to one-fourth of its earlier size. Effects at both ends are negligible. The gap between rotor and stator is separated into two areas. This is necessary to simulate a mechanical rotation. Every step in the transient simulation results in a new coupling with constraint equations between the nodes at the center lines. The rotating angle for every step can be calculated by using the electromagnetic torque and a simple torque equation.\nThree networks are created for the stator windings, the rotor windings and the slot wedges in addition to the FE-Model. The converter can be modelled as a voltage source with a variable amplitude and frequency. A DC voltage source and a rectifier are needed for the rotor, since the rotor will be coupled with its exciter during the starting process. The slot wedges are coupled to each other like a squirrel cage, but with the difference that there is no coupling from the last wedge to the first one.\nA particularity is the fact that the slot wedges are not continuously over the whole active length of the generator (cf. fig. II-A). Detailed studies on this subject show a raised resistance, which can be taken into consideration by a lower specific conductivity. An example can be seen in Table I.\nTwo special boundary conditions were chosen for the thermal simulation. On the one hand there will be an ideal thermal conductivity over the blue marked elements and on the other hand the blue marked elements will be replaced by air, similar to the electro-mechanical model. This simulates little gaps between the rotor iron and the slot wedges and shows the influence on the temperature rise. This is a worst case scenario, because the real existing gaps are much smaller as the blue marked elements in the thermal simulation. The green marked parts have ideal thermal conductivity over the whole simulation. It is assumed that the centrifugal force is high enough to put the wedges in radial direction.\nOnly the rotor is modelled for the thermal simulation. Heat transfer by radiation is ignored. Convection through the air gap of the generator can be taken into account with the equation 1, but most of the heat will be stored adiabatic.\nPtherm\nA = \u03b1 \u00b7\u0394\u03b8 (1)\nThe coefficient \u03b1 can be approximated by\n\u03b1 \u2248 8v2/3 [ W\nm2 \u00b7K\n], (2)\nwhere the parameter v is the air speed along the surface.\nTo get a closer look at the topic it is possible to create a torque curve for the asynchronous operation. This curve shows the torque over slip characteristics and is measured with a constant phase current amplitude of 4000 amps. Figure 4 shows the curve for this generator. If the slip turns over 7 per mille it is not possible to get an additional torque that is high enough to keep the generator in the synchronous operation. Because this range is really short, the generator is vulnerable to loose its synchronisation during the start-up process.",
            "In the example provided, called part two in table II, the rated frequency of 50 Hertz will be reached after a period of about 85 seconds. When the converter reaches this frequency, the voltage amplitudes will have 15 percent of the rated amplitudes. The exciter DC voltage will be set to 15 percent of the excitation voltage at no load over the whole simulation. To set up a field current, the system need to have a low speed of revolution. The real system has a little engine coupled with the shaft that can provide a starting frequency of 0.03 per unit.\nThe simulations shows that the generator follows the converter for about 15 seconds. After this duration the magnetic wheel reaches the critical angle of 180 degrees and the torque starts oscillating with the slip frequency. A comparison with Figure 4 shows the influence of the low slip range. The generator runs directly into the asynchronous operation, when there is a visible difference between the two curves in figure 5. The Fourier transformation shows three components, like mentioned in [1]. Two parts are at the slip and double slip frequency. The third part is a DC component which causes the little acceleration of the generator.\nThe currents in the slot wedges starts oscillating with the slip frequency, too. Figure III shows the currents in wedge one and wedge ten. The field in the spaces between the poles\nhas to be provided by the outer wedges. This causes higher currents and higher losses in the outer bars. Their amplitudes first rise with the slip frequency, until they reach a final value. The joule heating corresponds with the equation\nPloss = I 2 \u00b7R, (3)\nwhere R is the DC resistance. It is not possible to calculate the right DC resistance of a slot wedge, like mentioned in II-A. Therefore is a need of a comparison of the maximum losses between two different cases, with a good conductivity without gaps and a double resistivity. Figure III shows the two examples. The difference between the two curves increases with a rising stator current. While the resistance with a phase current amplitude of 4000 amps is negligible, it is not at higher phase currents. So it is necessary to use the losses corresponding with the higher resistance in the thermal simulation. The losses are stored for every time step and every element of the conductors in the model. This information can be used in the thermal simulation."
        ]
    },
    {
        "image_filename": "designv11_30_0001743_iccas.2015.7364719-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001743_iccas.2015.7364719-Figure2-1.png",
        "caption": "Fig. 2 Structure ofHAUV from the side.",
        "texts": [
            " HAUV must have the function of attitude and position control on the spot and the function of way-point movement in order to perform detailed examination and works [5]. PID controller and fuzzy controller are utilized for a certain level of control when there isn't a dynamical model about that underwater robot and its environment [6-8]. In this paper, way-point chasing simulation using PID controller as Matlab/Simulink is covered for tracking HAUV way-points and attitude control which are now being developed in our lab like Fig. 1. HAUV employed in this paper is what is reengineered existing HAUV [5]. HAUV has, like Fig. 2, the structure of aluminum frame, and it consists of 4 apparatus parts and 5 thrusters with skin on the outside. Apparatus parts are comprised of main controller part to control HAUV's movement, RTU(rotating thruster unit) to control the thrust and angle of the thruster, and battery part to supply power to body electronic system. Main controller part is composed of main PC, the controller for navigation, and sensor processing part, and RTU is made up of motor controller. Thrusters are composed of 2 omnidirectional tunnel thrusters, 2 horizontal tunnel thrusters, and 1 pitch tunnel thruster"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure5-1.png",
        "caption": "Figure 5. Velocity vector contours for standard dimensions.",
        "texts": [],
        "surrounding_texts": [
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nFigures 2, 3 show a finite-volume partition of the internal air domain of a field machine.\nTo obtain the distribution fields of flow rates in the body of the field machine, a series of calculations was carried out at different geometric parameters and drum rotation speeds.\nFigures 4 - 8 show the contours of the velocity fields and the velocity vectors for the initial geometric parameters (Figure 1). The rotation speeds of the drum and beater are 640 rpm and 2100 rpm, respectively. The maximum design flow velocity was 36.2 m/s. On the left side of the drum, an",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060\nincrease in the flow rate is observed, due to the narrowing of the space between the drum and the housing wall.\nA slight discrepancy in the display of velocities in the graphs of displaying fields and vectors is due\nto the peculiarities of displaying and visualizing numerical results.",
            "ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060\nIOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0001104_s10514-015-9420-9-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001104_s10514-015-9420-9-Figure5-1.png",
        "caption": "Fig. 5 a Quadruped model in simulation.The position and orientation of coordinate frame 1, attached to the torso, are used as the operational-space task to achieve stable trot control. b Connectivity tree for quadruped. Each leg is comprised of hip, thigh, and shank bodies that are each preceded by a 1-DOF abduction/adduction, hip swing, and knee joint, respectively",
        "texts": [
            " This section details the application of the recursive algorithms presented to the control of a quadruped trot. Diagonal leg pairs are synchronized during the trot, undergoing intermittent periods of stance during which leg forces are generated to propel and stabilize the torso, followed by the flight phase during which legs swing forward in preparation for touchdown again. The articulated legs are modeled with a series spring-actuation system acting around the knee. The quadruped is shown in simulation in Fig. 5a, with complete details of the system and its dynamic simulator in Palmer and Orin (2010). The quadruped weighs 76 kg in total and stands 60 cm high with the knees in a slightly bent configuration. The torso has a width of 35 cm and length of 1.2 m. The connectivity tree for the quadruped is shown in Fig. 5b, where each leg has three bodies: a hip, thigh and shank. Each of these bodies is preceded by a single-DoF leg abduction/adduction (ab/ad) joint, hip swing joint and knee joint, respectively. The system was simulated in the RobotBuilder environment developed by Rodenbaugh (2003). Contact forces were determined using a spring-damper penalty-based model of contact. Tangential forces were limited by static and kinetic friction coefficients of 0.75 and 0.6, respectively, where any friction violation led to slipping in simulation",
            " The work below assumes no such ground force sensor, but a similar approach can be taken if contact forces are known through the use of the original algorithm, and modification of Eq. 49 to t x\u0308d + bt = J\u0304 T t ST sw\u03c4 sw + J\u0304 T t ST st\u03c4 st . (50) The desired acceleration x\u0308d during stance is computed to achieve roll and pitch stability through proportionalderivative feedback but without affecting the forward, lateral, vertical and yaw motions predicted by the passive dynamics. The torso roll (\u03b3 ) and pitch (\u03b2) axes correspond to the x and y axes, respectively in the torso coordinate frame (shown in Fig. 5). Their desired accelerations are set as \u03c9\u0307x 1,d = k\u03b3 (\u03b3d \u2212 \u03b3 ) + k\u03b3\u0307 (\u03b3\u0307d \u2212 \u03b3\u0307 ), and (51) \u03c9\u0307 y 1,d = k\u03b2(\u03b2d \u2212 \u03b2) + k\u03b2\u0307 (\u03b2\u0307d \u2212 \u03b2\u0307), (52) where k\u03b3 , k\u03b3\u0307 , k\u03b2 , and k\u03b2\u0307 are control gains. Desired quantities \u03b3d , \u03b2d , \u03b3\u0307d , and \u03b2\u0307d are computed from cubic spline trajectories on the roll and pitch angles. These splines are initialized at the beginning of stance to match the pitch and roll setpoints selected at TOF. Desired accelerations for the remaining components seek to replicate the passive dynamics of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure14-1.png",
        "caption": "Fig. 14. Schematic of optimized recirculating hollow-shaft cooling structure with blades.",
        "texts": [
            " In addition, the pressure drop of the system is greatly increased due to the reduction of the inner diameter of tube. Obviously, from the overall consideration, reducing the inner diameter of tube is not a good optimization method. 3515 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on December 19,2020 at 20:47:14 UTC from IEEE Xplore. Restrictions apply. A recirculating hollow-shaft structure with blades at the bottom is proposed in this paper. Its specific structure is shown in Fig. 14. There are 6 blades distributed equidistantly in the circumferential direction in the bottom area. By comparing Fig. 15 and Fig. 6(b), it can be found that the flow of the recirculating hollow-shaft structure in the blade area is more turbulent than before, which will greatly increase the CHTC on the surface of the blade area. The results shown in Fig. 16 well verify the above analysis. The maximum CHTC at the corner of the blade, the local CHTC of the blade area can reach 2038W/m2/K, and the average CHTC of the entire heat exchange surface is 670W/m2/K, which is 49% higher than before optimization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000083_ssd.2019.8893176-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000083_ssd.2019.8893176-Figure3-1.png",
        "caption": "Fig 3. Distances between robot and obstacle",
        "texts": [
            " OBSTACLE AVOIDANCE CONTROLLER The obstacle avoidance is a basic requirement for mobile robots. Based on the acquired information from its sensors, the robot should move in any environments without collisions. In this paper, a Sugeno fuzzy logic controller is used to overcome the problem of encountered obstacles. The Khepera II robot is equipped by 8 infrared sensors of proximity which can detect obstacles. Only the 3 robot frontal sensors are used here. Each sensors can reflects a signal which intersect the obstacle to obtain 3 distances (see figure 3): \u2022 :Frontal distance, \u2022 : Left distance and \u2022 : Right distance. The Sugeno fuzzy controller have 3 inputs , and . Each distance , or is described by the following three fuzzy sets , , . The outputs of the Sugeno fuzzy logic controller are the left and right velocities and . The fuzzy logic controller of obstacle avoidance rules is based on the human knowledge and the relation between the sensor\u2019s information and the robot velocities. An example of a rule-base is given as follow: 978-1-7281-1820-8/19/$31",
            " It is noticeable that the best result of the travelled distance is obtained based on our proposed fractional order PI\u03bb controller. It is obvious that the distance between start point and desired point is decreased by 11% using the fractional order PI controller while comparing it by fuzzy logic controller [13]. 978-1-7281-1820-8/19/$31.00 \u00a92019 IEEE 81 2. Robot obstacle avoidance: The robot sensors send signals which intersect the obstacle to give birth to the distances between robot and obstacle , and (see figure3). If the obstacle is ahead the robot, so, at least one sensor can discover it. The obstacle avoidance by fuzzy logic controller combined with fractional order PI\u03bb one is depicted in following picture (figure 6). The desired robot target position is firstly (300 mm, 300mm) and secondly is (400 mm, 400mm) and in both situations the start point is placed at (0, 0) with initial orientation of 0\u00b0. The left and right velocities are presented in figures 7 and 8: The simulations results demonstrate that when the Khepera II robot discovers a close obstacle by its frontal sensor, it will spontaneously produces the adequate left and right speeds (see figures 7,8) using the Sugeno fuzzy logic obstacle avoidance controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure4-1.png",
        "caption": "Fig. 4. Spatial involutoid for hypoid gearing.",
        "texts": [
            " Expressing the gear ratio g in terms of the evolutoids gives g \u2261 \u03c9 o \u03c9 i = r bi r bo \u221a 1 \u2212 ( C li \u00b7 b i ) 2 \u221a 1 \u2212 ( C lo \u00b7 b o ) 2 = r bi r bo sin \u03d1 i sin \u03d1 o (9) where r bi Radius of input evolutoid (input base circle) r bo Radius of output evolutoid (output base circle) \u03d1 i Angle of incidence with input evolutoid \u03d1 o Angle of incidence with output evolutoid C li Unit direction of contact normal (input gear) C lo Unit direction of contact normal (output gear) b i Binormal of evolutoid (input gear) b o Binormal of evolutoid (output gear). For spur cylindrical gears, C l \u00b7C i = 0 and the gear ratio is also equal to the ratio of radii of the base circles. Depicted in Fig. 4 is a transverse curve between the toe and heel of a hyperboloidal pitch surface. Also presented is the evolutoid defined by the transverse curve with constant pressure angle \u03c6 and spiral angle \u03c8 . Coordinates for are based on a pitch point t p on the transverse curve. Associated with t p is the reference evolutoid point E re f and reference trihedron \u3008 t re f , n re f , b re f \u3009 . Cartesian coordinates for the involutoid are defined by a constant direction in the rectifying plane of the evolutoid as the taut chord is unwrapped: = \u23a1 \u23a3 cos ( \u03b8 \u2212 \u03b8re f ) \u2212sin ( \u03b8 \u2212 \u03b8re f ) 0 sin ( \u03b8 \u2212 \u03b8re f ) cos ( \u03b8 \u2212 \u03b8re f ) 0 0 0 1 \u23a4 \u23a6 [ E re f + ( r b \u03b8sin \u03d1 ) C l ] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002183_2168-9792.1000163-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002183_2168-9792.1000163-Figure3-1.png",
        "caption": "Figure 3: Quadcopter Virtual model.",
        "texts": [
            " The stiffness of the solid structure is lesser than the hollow structure. The torsional stiffness of a closed square cross-section is greater than the closed circular section. Therefore closed square cross sectional hollow frame is used. This reduces overall weight. The stiffness can be varied by changing crosssectional profile dimensions and wall thickness. Therefore box type frame is chosen (Figures 1 and 2). The quadcopter virtual model with motors and propellers assembly is designed using CATIA software as shown in Figure 3. The parameters of the quadcopter frame are shown in Table 1. The propulsion system consists of motors, propellers, Electronic Speed Controllers (ESCs), batteries and propellers. Both the motor and propeller combination produces thrust and moves the vehicle upwards. As the estimated all up weight is considered to be 1.5 kg, the thrust requirement from four motors should be double that of 1.5 kg. Therefore each motor should be able to produce 850g of thrust force. Motors are selected based on their Kv rating"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000682_demped.2015.7303706-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000682_demped.2015.7303706-Figure1-1.png",
        "caption": "Fig. 1: Positioning of accelerometers on generator bearings. The accelerometers are installed at the load zone for more efficient and consistent fault detection.",
        "texts": [
            " The mechanism of bearing creep and its development are described in section III. The tools utilized towards the detection of inner race bearing creep are presented in section IV. Fault detection of two different bearing creep severity cases is shown in section V. Finally, sections VI and VII presents the discussion and conclusions of this work respectively. Vibration analysis has been the most wide-spread condition monitoring technique applied on wind turbines generators employing accelerometers installed radially at the load zone, as shown in Fig. 1 [10]. Tracking of speed related spectral components describing the shaft dynamics, such as the first and higher orders running speeds, is usually implemented in condition monitoring systems along with broadband measurements in various frequency ranges for overall vibration evaluation and early stage bearing defects [7]. Alternatively, an envelope can be set over the considered healthy vibration signature, where an alarm is triggered when a frequency component exceeds the above mentioned predefined limit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000746_1.4028998-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000746_1.4028998-Figure2-1.png",
        "caption": "Fig. 2 Schematic of contact between the sum surface and a rigid flat plane under the action of both normal and shear forces",
        "texts": [
            " The expected value of a quantity of interest X(n, g) for a random asperity is then calculated as X \u00bc \u00f0\u00f0 X\u00f0n; g\u00def \u00f0n; g\u00dedndg (12) The expected macroscopic value of the quantity of interest is calculated as hXi \u00bc Dp \u00f0\u00f0 X\u00f0n; g\u00def \u00f0n; g\u00dedndg (13) where Dp is the density of asperities defined in Eq. (5). This process of calculating the expected macroscopic value of the quantity of interest from its microscopic value is called statistical homogenization. For the contact between a sum surface and a rigid flat plane under the action of both normal and shear forces shown in Fig. 2, the macroscopic estimates can be calculated by considering the response of a single asperity. Consider an asperity of normalized height n and curvature g. Let the normal force needed to deform the asperity by a normalized distance h\u00bc d/r be P*(n, g), keeping in mind that the deformation may be in the plastic domain. The normalization constant r is the RMS roughness of the surface introduced earlier in Eq. (3). Let Q*(n, g) be the tangential force the asperity is capable of supporting before yielding reaches the surface of the asperity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure2-1.png",
        "caption": "Figure 2. Wheel and rail cross-section profiles and x-y plane coordinate system.",
        "texts": [
            " Furthermore, an example calculation is performed, the results are analyzed, and a series of conclusions are obtained. Structure of wheel/rail profiles and the principles behind determining contact positions Figure 1 shows a simplified schematic drawing of wheel/rail contact on straight track, where d is the distance between the center lines of the left and right rail cross-sections, T is the distance between the backs of the two wheel flanges and W is the axle load. When a train runs on the track, the relative position between the wheel and rail is not fixed along the x-direction (Figure 2). For example, when running on curved tracks, the whole train will move right or left under the effect of centrifugal force; thus, the relative position of the wheel and rail will change and this causes the wheel/rail contact position to change in response. Therefore, the profile geometry parameters at different points on the wheel and rail are not the same, and thus the wheel/rail contact stress will change as the contact position changes. In this paper, a single wheel and rail are taken as a unit for study. In order to study wheel/rail contact stress, a method must be established to determine the center position of contact points in different contact situations; however, the fact that the relative position of the wheel and rail is not fixed makes it difficult to find the center. Figure 2 shows the cross-section profiles of the wheel and rail; they are composed of arcs and lines. Points A, B, D, E, F, H, I, J and K are dividing points of the arcs or straight lines on the wheel\u2019s surface profile, and point C is the vertex of the wheel flange. Points A1, B1, C1, D1, E1 and F1 are dividing points on the rail\u2019s surface profile and point G is the vertex of the rail profile. The coordinate system of the x-y plane is the same as the one plotted for the wheel tread in China\u2019s National Standard for Passenger Train Wheels and Rails"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure9-1.png",
        "caption": "Fig. 9. Substituted RUSR mechanism. (a) Joint centers K and Ko are arbitrarily chosen for velocity analysis, (b) Kconj is chosen and conjugate center Kconj o is located at a corresponding distance \u03c1 for acceleration analysis.",
        "texts": [
            " Using these coordinates, the two new substitution centers are chosen by means of Eq. (60). A value of z is chosen arbitrarily, then the corresponding zowill be determined from that equation. After the substitution, the acceleration analysis of themechanism is done using the existing methods for lower pair mechanisms. But it should be reminded that before this second step of acceleration analysis, velocity analysis of only the link-lmay need to be reworked since the joint centers of this link changed. Fig. 9 shows an illustration of the mechanism in Fig. 1 with substituted connection. The expression for zo in Eq. (60) does not seem to get rid of the twist coordinates for any choice of z expressed in terms of the curvatures of contacting surfaces. This means that the centers of the substitute-connection cannot be obtained using only the curvature values obtained by mere examination of the profiles of the contacting surfaces in a general scenario. But it can be easily shown that substitution using geometry parameters alone is possible for local spherical surfaces in contact, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003733_icems50442.2020.9290866-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003733_icems50442.2020.9290866-Figure6-1.png",
        "caption": "Fig. 6. Rotor assy",
        "texts": [
            " Since flux range of the proposed motor is correspond to the maximum and minimum field current condition, it can be seen that characteristic of copper loss reduction can be estimated by using analysis results of both conditions. From Fig. 5, about 27.3% of weakening current can be reduced by using the proposed motor at 20000 r/min condition. From this result, up to 47.1% copper loss reduction can be obtained by using the proposed motor. 1) Prototype motor In this term, validation of the simulation results shown in Fig. 5 is confirmed using experimental results. The motor specifications are same as shown in Table II. In Fig. 6, prototype rotor assy is shown. From Fig. 6(a), insert process of coil winding unit can be conducted to fill the gap between laminated rotor core and small size core. On the other hand, it can be assumed that coil end and crossing line of the field coil are affected by shearing stress caused by rotary centrifugal force. To avoid the disconnection trouble in the field coil, resin mold is given as shown in Fig. 6(b). The resin material is selected as heat curing type. 2) Motor bench Fig. 7 shows a motor bench. Torque is measured using a torque meter. In the load testing, rotational speed is set by the load motor. PE-Inverter and PE-Expert made by Myway Company were selected as the inverters and drive controller. Limit phase current and voltage were set to 500Arms per phase and 48V-DC. - 726 - Authorized licensed use limited to: CALIFORNIA INSTITUTE OF TECHNOLOGY. Downloaded on August 13,2021 at 07:27:59 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002927_tec.2020.2983187-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002927_tec.2020.2983187-Figure4-1.png",
        "caption": "Fig. 4. Schematic diagram of the \u25b3-connect three-phase windings.",
        "texts": [
            " The claw pole alternator has a compact structure. In order to improve the power density as much as possible, the stator is required to have a high coil fill factor, that is, the diameter of the copper wire should not be too large. On the other hand, in order to increase the output current, the stator coil resistance is required to be as small as possible. Therefore, the common practice is to first connect two or three wires with smaller diameter in parallel as one wire, and then wrap them around the stator teeth. Fig. 4 shows the schematic diagram of the \u25b3- connect three-phase windings, in which each phase winding consists of two parallel wires. In other words, Ab1-Ac1 and Ab2-Ac2 shown in Fig. 4 are two wires of phase A. The two wires are connected in parallel as one new \u201cwire\u201d and then this new \u201cwire\u201d is wound on the stator teeth in the way shown in Fig. 3. Phase B and phase C are similar to phase A. Thus the three phases totally have six wire ends and the six wire ends are connected as \u25b3-connect. The \u25b3-connect three-phase windings have three terminals which would be connected to the three-phase diode rectifier bridge. Since each phase winding has two parallel wires, the stator winding interturn short circuit fault of the same phase winding would include the following two situations: short circuit between turns of the same wire (see in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003291_012008-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003291_012008-Figure2-1.png",
        "caption": "Figure 2. Diagram of forces acting on active rear axles of multi-purpose forest fire fighting vehicle",
        "texts": [],
        "surrounding_texts": [
            "In Vietnam, most durability tests are carried out on test bench with static load modes. In this way, they are limited in assessing the durability of vehicles in actual operating conditions. Therefore, they need to use dynamic loads as input parameters for durability determination models that have increasingly been developed in recent years. The chassis is a girder system bearing loads installed and loaded, receiving and transmitting jets during vehicle operation. Besides, it is influenced by vibrations from engines, powertrain, etc. A study of the load mode applied on the chassis is only within atheoretical model but there are no practical tests to assess the reliability of this model [1, 2, and 3]. Stemming from that reason, the author proposed a method to determine the dynamic load from the road surface to the cement frame by an experimental method with existing equipment in Vietnam. The purpose of the experiment is to measure the pavement reaction force acting on the sphere when the vehicle passes the standard scale. The reaction value from the road surface to the axle via suspension will affect the chassis. Experimental results are compared with the theoretical results to verify the simulation model 10th TSME-International Conference on Mechanical Engineering (TSME-ICoME 2019) IOP Conf. Series: Materials Science and Engineering 886 (2020) 012008 IOP Publishing doi:10.1088/1757-899X/886/1/012008 to determine the dynamic load, as the reason for the calculation of durability of a multi-purpose forest fire truck. Calculations that determine the impact on the chassis are carried out on multi-purpose forest fire fighting vehicle manufactured by Vietnam. This is a vehicle integrating many forest fire fighting functions including cutting trees, cleaning garbage grass, opening roads to create a fire isolation corridor; fire sprinkler with wide spray area; create high-pressure wind spray on the fire; using sandy soil on spot to extinguish the fire has been designed based on the vehicle URAL 4320, which has three active main axles used in the military field. Studies to determine the force reaction from the road surface to the rear axle are limited, especially the method of experimentally determining when the vehicle is moving on the road. Many studies on the durability of rear axle cover under the effect of the effective loads [7, 8, 9, and 10]; these studies indicate the stress and deformation that appear on the active bridge cover. In their research, M H Trinh and colleagues [11] have determined the dynamic load from the road acting on the rear axle of the truck by experiment method, tested on real roads, vehicles passing through sinusoidal models when the vehicle is moving at different speeds. 2. Model of load determination on the bridge shell As can be seen in the diagram, the shell is subject to the reaction of the jet from the road surface Fz31, Fz32 (N) and force from leaf spring Fzn31 v\u00e0 Fzn32 (N); a shows the distance from the point of 10th TSME-International Conference on Mechanical Engineering (TSME-ICoME 2019) IOP Conf. Series: Materials Science and Engineering 886 (2020) 012008 IOP Publishing doi:10.1088/1757-899X/886/1/012008 the jet from the road surface to the point of the force from leaf spring (m); b displays the distance from two points of the jet from the left and right pavement (m)."
        ]
    },
    {
        "image_filename": "designv11_30_0002504_iceeot.2016.7755478-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002504_iceeot.2016.7755478-Figure1-1.png",
        "caption": "Fig. 1. Machine fault simulator.",
        "texts": [
            " The MRA consists of the application of the DWT based on the decomposition of a signal using a bank of lters, but not able to decompose half superior frequency band of a signal to resolve this problem nowadays, the Wavelet Packets Transform (WPT) constitutes the last improvement of the DWT, is replacing the MRA. Though the wavelet coefficients can be used directly as features, WPT also have Disadvantage; its application is a complex task due to the diversity of critical parameters that must be selected, such as the mother wavelet and the decomposition level, while there is not an established methodology to select them. II. EXPERIMENTAL SETUP The experimental setup consists of a Machine Fault Simulator (MFS) is shown in figure 1, having an induction motor coupled with a rotating shaft supported in the Rolling Element Bearing (REB), a CSI accelerometer, a portable device vibscanner to record vibration signal, a data acquisition card to collect current signal and a computer with MATLAB R2009a software. The details of specimen (REB) are listed in table I. III. DATA ACQUISITION REB faults are classified on the basis of their location as Inner Race Defect (IRD), Outer Race Defect (ORD) and Rolling Ball Defect (RBD). These defects produced artificially and couse their certain fault frequency to appear in the bearing vibrations as well as in motor current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001508_0954406214553983-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001508_0954406214553983-Figure6-1.png",
        "caption": "Figure 6. Collision of two rigid bodies.",
        "texts": [
            " The variation of the kinetic energy is given17 by the relation Ec \u00bc 1 2 v0 T g\u00bd MC\u00bd v0 1 2 vf gT g\u00bd MC\u00bd vf g \u00f040\u00de Since the matrix g\u00bd MC\u00bd is a symmetrical one, it results in MC\u00bd T g\u00bd \u00bc g\u00bd MC\u00bd \u00f041\u00de and the equality 2 Ec \u00fe 2 Ep \u00bc 2 v0 T g\u00bd MC\u00bd vf g v0 \u00f042\u00de Taking into account equations (34), (39), and (31), one deduces Ec \u00bc 1 2 P2g Pv0n \u00f043\u00de which is formally identical to equation (9). Collision of two rigid bodies To study the collision of two rigid bodies at point A, we consider the schema in Figure 6 and the previous notations adapted with indices 1 and 2. The drawn reference systems are those of the principal central axes of inertia. The column matrices Nj , j \u00bc 1, 2, are defined by the projections of the vectors n, CiA n, i \u00bc 1, 2, respectively, onto the axes of the systems Cixiyizi, i \u00bc 1, 2. Analogically, one defines the column matrices fv0j g and vj , j \u00bc 1, 2. It results that the inertances gj, j \u00bc 1, 2, are given by the relations gj \u00bc Nj T g\u00bd MjC 1 Nj \u00bc 1 mj \u00fe d2j Jjx \u00fe e2j Jjy \u00fe f2j Jjz , j \u00bc 1, 2 \u00f044\u00de From equations (33\u201335), (39), and (43), for j \u00bc 1, 2, one obtains the equalities MjC vj v0j n on o \u00bc 1\u00f0 \u00dejP Nj \u00f045\u00de vj v0j n o \u00bc 1\u00f0 \u00dejP MjC 1 Nj \u00f046\u00de vjn v0jn \u00bc 1\u00f0 \u00de jPgj \u00f047\u00de Ejp \u00bc 1 2 P2gj \u00f048\u00de at University of Sydney on October 7, 2015pic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003680_icee50131.2020.9260982-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003680_icee50131.2020.9260982-Figure1-1.png",
        "caption": "Fig. 1. The schematic view of the cable of CDPR.",
        "texts": [
            " The practical workspace of the robot is a rectangular cube with dimensions of 360cm \u00d7 220cm \u00d7 160cm. Table I refers to the workspace, which is considered for training the neural network. The manipulator consists of four cables and an EE. The EE is a square with 10 cm sides. A coupling, serving as a mechanical fuse, connects the actuator to the rolling system. Further information regarding this CDPR can be found in [12]. The solution of IKP for CDPRs has only one solution. However, in the case that the cables are not fully in tension, the acquired solution is not applicable. As shown in Fig. 1, and denote the connection point of the cable to the actuator and EE respectively; represents Authorized licensed use limited to: McMaster University. Downloaded on May 25,2021 at 13:33:32 UTC from IEEE Xplore. Restrictions apply. the cable length; and are constant vectors in the global and local coordinate systems, respectively. Center of the mentioned coordinate systems is denoted by O and E, respectively. Also, point E is the EE center of mass. Therefore, the loop-closure position equations are written as the following equation: = + \u2212 (1) where shows the position vector of the robot in the global coordinate, and R represents the rotation matrix from the origin of the local coordinate system to the origin of the global one [13]: = \u2212 ++ \u2212\u2212 (2) In which c and s stand for the cosine and sine of the indexed argument, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002813_ilt-06-2019-0239-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002813_ilt-06-2019-0239-Figure1-1.png",
        "caption": "Figure 1 Schematic diagram for fully textured journal bearing",
        "texts": [
            " (2019b) have used an ultra-short pulse laser to fabricate a single pocket, line-, cross- and dot-like texture. Based on the studies conducted in this field and previously published researches, we find that most of the papers have been interested in the effect of the shape of the dimple\u2019s bottom area and its dimensions on the bearing performance. However, the thermo-hydrodynamics of lubricant in a surface textured journal bearing was ill studied. The fully textured journal bearing geometry in 2D and 3D is shown in Figure 1. The 3D axis is represented in a circumferential direction with length 2pR, a longitudinal direction with length L and a radial direction with length h in the x, z and y-directions, respectively. The radial height for smooth bearing is h; however, it increases by6ht for a textured surface. In this study, a group of 24 different shapes of dimples has been studied, half of them has a convex shape which decreases Effect of liner surface texture AhmedM. Saleh et al. Industrial Lubrication and Tribology Volume 72 \u00b7 Number 3 \u00b7 2020 \u00b7 405\u2013414 the lubricant film thickness ( ve h) as shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure19-1.png",
        "caption": "Fig. 19. Shape rolling of a long product (",
        "texts": [
            " It shows that even though smoothing with consistent normals provides an enhanced description of contact area [24], the use of normal voting gives an even better description of tool geometry and consequently of contact interactions during the entire simulation. Fig. 18 also shows that equivalent strains are more homogenous and have lower values: the maximum of equivalent strain is about 0.64 using consistent normals and about 0.60 using normal voting. This example shows the importance of the accurate description of contact and of enhanced smoothing procedure. The geometries of tool and bar employed in the shape rolling simulation are shown in Fig. 19 \u2013 right. The material is viscoplastic (Norton Hoff law (3) with: K = 158 MPA s0.18, m = 0.18) and sliding contact. The mesh size is approximately h = 9 mm and the tool mesh size is about h = 9.8 mm. For this industrial process, smoothing significantly improves contact prediction because the contact area is smoother in space and more regular in time, as can be seen in Fig. 20. The equivalent strains (Fig. 21) are more homogenous and lower; the maximum value decreases from 1.96 to 1.82, which corresponds to expecta- tions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003484_0954406220963148-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003484_0954406220963148-Figure15-1.png",
        "caption": "Figure 15. Experiments with inspection robot.",
        "texts": [
            " In summary, the static global path planning algorithm based on the improved bat algorithm has a stronger global search ability, which effectively prevents the algorithm from falling into a local optimum and has the ability to skip the local optimum. The local search ability is improved, and the planned path is shorter with fewer turning points and better quality. In order to verify the effectiveness of the improved algorithm in this paper, some experiments are carried out under different obstacle scenarios indoors using a differentially driven mobile robot equipped with ROS system. Mobile robot hardware includes: wheel motor, motor driver, embedded PC, lithium battery, lidar and IMU, as shown in Figure 15. Four indoor environments are designed in this experiment, including barrier-free environment, static obstacle environment, environment with rightangle bends and narrow passages, and environment with U-shaped bends, to test the effect of the improved path planning algorithm. This experiment is completed in a 300 360 indoor barrier-free environment (Figure 16(a)), and a good environment map is built as Figure 16(b) through a lidar. In rviz, the target point pose of the robot is set through the green arrow in Figure 16(c), and the system calls the global path planning algorithm to generate the optimal path, shown as the red line in Figure 16(d)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure12-1.png",
        "caption": "Fig. 12. Schematic of optimized recirculating hollow-shaft cooling structure for reducing inner diameter of tube.",
        "texts": [
            " However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization. Although changing the bottom to a cone-shape can improve the CHTC of the bottom area, the overall CHTC does not increase much. A simple optimization method is to gradually reduce the inner diameter of the bottom stationary cooling tube, so that it enhances the disorder of the bottom fluid, thereby improving the value of the overall average CHTC. The detailed structure is depicted in Fig. 12. Fig. 13 shows that the maximum value of the CHTC of the structure after optimization is 1.57 times that before optimization, and the average CHTC is increased by 16%. However, the friction loss of the wall surface increased by 78%. In addition, the pressure drop of the system is greatly increased due to the reduction of the inner diameter of tube. Obviously, from the overall consideration, reducing the inner diameter of tube is not a good optimization method. 3515 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002769_ls.1499-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002769_ls.1499-Figure5-1.png",
        "caption": "FIGURE 5 Different adapting brackets and driving elements to test the different combinations of rolling bearing and lubrication method",
        "texts": [
            " At this point, it should be apparent that in Figures 2 and 4A, the assembly is shown with a thrust roller bearing (81107 TN), but in Figures 3 and 4B the new assembly is shown with a tapered roller bearing (32008 X/Q). In fact, with the proposed assembly rolling bearings with dimensions up to the dimensions of the 32008 X/Q tapered roller bearing can be tested. Table 1 shows the types and dimensions of the rolling bearings that can be tested and Table 2 shows the possible combinations of speed and load. In order to test the different rolling bearings presented in Table 1, different adapters must be used. Figure 5 shows the different options for testing the different rolling bearings either oil or grease lubricated. In Table 3, the rolling bearings that can be tested are shown along with what the nominal operating conditions represent relative to the dynamic basic load rating and limiting speed. According to SKF7 for grease lubricated rolling bearings, the free volume of grease to apply in order to have adequate lubrication is a function of a speed factor A = n dm. Table 4 shows the initial grease fill % (calculated according to SKF7) as well as the volume of oil needed for the possible rolling bearing tests"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000608_jae-141776-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000608_jae-141776-Figure1-1.png",
        "caption": "Fig. 1. PM rings: Z-axis \u2013 the axis of symmetry, rin-inner radius, rout-outer radius and h-height of the rings. (a) Magnetic polarization (J) is axial. (b) Magnetic polarization (J) is radial.",
        "texts": [
            " However the researchers have not addressed the effect of axial, radial and angular displacements of the rotor magnet ring on the generated magnetic field. In the present paper, the magnetic field created in PM bearings is analysed by considering the three translational (x, y and z) and two angular (\u03be and \u03b3) degrees of freedom of the rotor magnet ring using Coulombian model and vector approach. In addition, variations of magnetic field components in the air gap due to movements (three translational and two angular) of the stator ring magnet are calculated and presented. The PM rings with axial and radial magnetization are shown in Fig. 1. The magnetic field created by the PM ring is evaluated using the Coulombain approach. In this approach, the PM is represented by the surfaces with fictitious magnetic pole surface densities in the direction of polarization. A and B the fictitious charged surfaces of the magnet rings. Let us consider the elements A1 and B1 on the surfaces A and B of the magnet rings as shown in Fig. 1. The magnetic field created by the surface element A1 of the surface A at any observation point P (x, y, z) of the space is given by the Eq. (1). HA1 = (+J)SA1 4\u03c0\u03bc0r3A1P rA1P (1) where J is magnetic surface flux density, SA1 is the surface area of the element A1 and rA1P is the distance vector between the element A1 and the point P . Similarly, the magnetic field created by the element B1 of the surface B at the same observation point can be written as: HB1 = (\u2212J)SB1 4\u03c0\u03bc0r 3 B1P rB1P (2) where SB1 is the surface area of the element B1 and rB1P is the distance vector between the element B1 and the point P"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003291_012008-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003291_012008-Figure4-1.png",
        "caption": "Figure 4. Diagram of the normal force",
        "texts": [
            " The objective of the experiment is to measure the dynamic load from the road surface acting on the bridge in the process of passing the standard scale. Experimental results are compared with calculation results by dynamic models to verify theoretical models. The parameter to be measured in the experiment is the force exerted from tires onto the rear axle in a vertical direction Fz (N) The method of measurement determines the forces applied from tires to the rear axle vertically with the application of the Tenzo switch with bending load on any beam. Figure 4 is a diagram showing the measurement structure system that determines the bending load on a beam. The measuring system includes sensors, which receive, amplify, convert, process, display and store measuring results. The electrical signal (current or voltage) from the sensors is amplified many times, then through the A/D converter into a digital signal transmitted to the software installed in the computer (PC). 10th TSME-International Conference on Mechanical Engineering (TSME-ICoME 2019) IOP Conf",
            "1088/1757-899X/886/1/012008 To measure the vertical legal force, the research team used a four-resistor device. Four tenzo resistors are pasted in parallel and arranged symmetrically on the upper and lower sides of the bridge at 12 o'clock and 6 o'clock positions (the upper surface of 02 leaves, the lower surface of 02). 1) Bridge beam; 2) Tenzo; 3) Amplifiers and A/D Converter; 4) Computer. Experimental devices include: a. Resistor bridge A diagram of the axle resistorand a measurement resistor bridge is shown in Figure 4 above. Tenzo resistors are applied directly to the outer surface of the axle. Before that, one must clean the shaft surface to achieve the required roughness, clean with gasoline and acetol. To measure the vertical force acting on the rear axle proactively, four Tenzo resistive tenzo leaves (120V) are affixed to the upper and lower surfaces of the axle, facing on two leaves, and facing below two parallel leaves and opposite. Four Tenzo are arranged passes through the active axle axis (Figure 5)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000816_melcon.2014.6820563-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000816_melcon.2014.6820563-Figure2-1.png",
        "caption": "Fig. 2. General-purpose multi-mass model of a driven mechanism.",
        "texts": [
            "00 \u00a92014 IEEE 373 that takes into consideration this imperfection of the mechanical part consists of introducing the multi-mass model. This new concept divides the mechanical part into subsystems. Each rotating part (resp. each part in translation) is represented by a mass with an inertia Ji and a velocity \u2126i (resp. by a mass with a weight Mi and a linear speed Vi). Moreover, these two consecutive rigid parts are connected by an elastic link with an elasticity factor ki,i+1 and a damping factor \u03b2i,i+1. Fig. 2 presents the modeling of an electrical drive system by a multi-mass system. Even though the multi mass model describes accurately the mechanical system, its complexity is a sort of weakness: there are a lot of differential equations that rule the behavior of the system. Consequently, complex numerical computations should be done in order to resolve them. Thus, the engineer should create a new model of the mechanical system that both reduces the order of its differential equations and describes it precisely: a threemass model or a two-mass model could be used, as illustrated respectively in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000229_saci46893.2019.9111599-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000229_saci46893.2019.9111599-Figure6-1.png",
        "caption": "Fig. 6. FBG displacement sensor for RAMIS, developed by Liu et al. [54].",
        "texts": [
            " FBG sensors are based on the strain-induced shift of Bragg wavelength; a particular wavelength of light that is reflected by the optical fiber, while all other wavelengths are transmitted [64]. These optical fiber sensors possess small physical size and are also tolerate the high temperature during autoclave sterilization, furthermore immune to electromagnetic interference. The mentioned features make FBG force sensors an optimal candidate for RAMIS palpation probes, as already integrated into RAMIS instruments by various research groups (Fig. 6) [53]\u2013[56], [59]. C. Palpation with vision-based force estimation The applied force during RAMIS can also be estimated based on the endoscopic camera images. The major benefit of this technique is that no additional device needs to be added to the operating room setup. Furthermore, bio-compatibility and sterilization are solved already in the case of RAMIS endoscopes. This technique involves several different types of image processing methods, such as feature extraction, filtering of light reflections, and also can be approached using neural networks, but the crucial parts are usually the reconstruction of the tissue surface and handling the inhomogenity in the tissue"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003029_s00170-020-05351-5-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003029_s00170-020-05351-5-Figure2-1.png",
        "caption": "Fig. 2 Meshes and boundary conditions",
        "texts": [
            " The Energy Conservation Equation was mainly used to discretize and solve the heat transfer mechanism. The discrete particles were tracked through the computational domain through solving the Lagrangian equations. The grids were generated in CFD-GEOM (version 2016.0.0.17). The geometry of the physical domain, as shown in Fig. 1, was sized at 700mm (length) \u00d7 1495mm (weight) \u00d7 500mm (height), including one blowing nozzle on the top and two suction tunnels at both sides of the working chamber. As shown in Fig. 2, structured grids were generated for the entire model apart from the irregular area such as the blowing nozzle and the corners in the suction devices, for the purpose of obtaining better convergence. The meshes along the Z-direction away from the working plane (Z \u2265 300 mm and Z \u2265 \u2212 300 mm) were gradually dispersed and reduced under the power law to avoid computational expanses. Based on the grid independence study carried out in our previous study [10], the numbers of grids of 1,710,000 have been proved to obtain the best match to the experimental results, which can be employed in this simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure14-1.png",
        "caption": "Fig. 14 Three-carrier planetary gear train\u2014multiplier of wind turbine. a Kinematic and structural schemes. b Determination of speed ratio. c Determination of efficiency",
        "texts": [
            " 10 shows a reverse gear train, carrying forward and backward. The Fig. 11 shows a change-gear (gearbox), carrying out two gear-ratio steps (speeds). Figure 12 shows three cases of using two-carrier compound planetary gear trains with a one compound shaft, but with four external single (i.e. not-compound) shafts. Figure 13 shows the structural schemes of three-shaft three-carrier and fourcarrier compound gear trains. The first are used as reducers or multipliers, i.e. with F = 1 degree of freedom, while the second\u2014as change-gears. Figure 14 shows three-carrier compound gear train that works as a multiplier in the powerful wind turbine (Giger and Arnaudov 2011). Figure 15 shows how to determine the load spectrum of an element of fourcarrier change-gear while working with different gear-ratio steps (speeds). Unlike the most commonly used methods of Willis and Kutzbach-Smirnov that use the angular and peripheral velocities, at the presented method here another parameter of mechanics is used\u2014the torque. This alternative method is characterized with the following: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002463_epe.2016.7695442-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002463_epe.2016.7695442-Figure1-1.png",
        "caption": "Fig. 1. Relationship between the -frame and the dqframe. The q-axis is defined as the direction of the electromotive force vector. The -axis is defined as the direction of the output voltage vector of the inverter.",
        "texts": [
            " It is noted that the MTPA control using the hill climbing method is not effective in the sensorless FOC because the sensorless FOC requires the motor parameters in the control strategy. In addition, the variation of the d-axis current by the hill climbing method disturbs the position estimation system. This paper will be organized as follows; first, the principles of the V/f control is introduced. Next, the MTPA control based on the hill climbing method is explained. Finally, the experimental results are shown to confirm the validity and the effectiveness of the proposed method. A. V/f control Fig. 1 shows the relationship between -frame and dq-frame on rotating frame which is synchronized to rotating speed of the motor. In the permanent-magnet synchronous motor control, the d-axis and the q-axis are generally defined as the direction of the flux vector in the permanent magnet and the electromotive force vector, respectively. Therefore, the identification of the flux vector is important to apply in the FOC on dq-frame. On the other hand, the -axis is defined as the direction of the output voltage vector of the inverter and the -axis is defined as the -axis delayed by 90 degrees"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002927_tec.2020.2983187-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002927_tec.2020.2983187-Figure9-1.png",
        "caption": "Fig. 9. Finite element model of the magnetic field simulation under the stator winding interturn short circuit fault.",
        "texts": [
            " However, when the stator winding interturn short circuit fault occurs, the current in the same conductor before and after the short circuit point is different. For the phase winding that consists of two parallel wires, there would be four different currents (I1, I2, I3, I4) in each stator slot under the stator winding interturn short circuit fault, as shown in Fig. 5, so four conductors need to be modelled in each stator slot. On the other hand, the interturn short circuit fault destroys the periodicity of the alternator, so the whole machine model must be used for simulation instead of the one sixth model (see in Fig. 9). As a generator, the field current and rotor speed of the claw pole alternator are the two parameters that need to be set for the simulation of the magnetic field and magnetic force. The phase current is also the simulation result in the alternator, just like the magnetic field. On the contrary, the phase current is a main parameter that needs to be set before the simulation for electric motors. When the stator winding interturn short circuit fault occurs in the claw pole alternator, the current in the short circuit phase becomes complicated that there are multiple currents with different amplitudes and phase angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure5-1.png",
        "caption": "Fig. 5. A car-like wheeled mobile manipulator with a two-link robot arm.",
        "texts": [
            " Virtual Joint Method for Kinematic Modeling of Wheeled Mobile Manipulators 1063 where zi\u20131 is the unit vector of the i-th joint axis of the transformed system, ree is a vector from Op to On, and ri\u20131 is a vector from Op to the origin of a coordinate frame at the i-th joint, following the standard Denavit-Hartenberg convention [14]; all these vector quantities are expressed with respect to frame \u03a30. 3.3. Case example We choose a simple planar two-link robot arm asymmetrically mounted on a car-like mobile platform as shown in Fig. 5(a). Applying the virtual joint method to the mobile manipulator, we obtain an equivalent 4-DOF articulated manipulator as shown in Fig. 5(b). The geometric Jacobian for this equivalent system is formulated as follows: 0 0 1 2 2 3 3 0 2 3 ( ) ( ) ( ) , 0 v ee ee ee = \u00d7 \u2212 \u00d7 \u2212 \u00d7 \u2212\u23a1 \u23a4 \u23a2 \u23a5 \u23a3 \u23a6 J z r r z z r r z r r z z z where 1 [0 01] for 0,2,3, [cos sin 0] , T i T i \u03c6 \u03c6 = = = z z 0 0 0 2 0 0 1 01 1 01 2 012 3 1 01 1 01 2 012 , , 00 0 , , 0 0 p pm p m p m m m ee m x x dc hsx y y y ds hc x L c x L c L c y L s y L s L s + \u2212\u23a1 \u23a4 \u23a1 \u23a4\u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5= = = + +\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u23a3 \u23a6\u23a3 \u23a6 \u23a3 \u23a6 + + +\u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5= + = + +\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6 r r r r and 0 cos ,c \u03c6= 01 1 cos( ),c \u03c6 \u03b8= + 012 1 2 cos( ),c \u03c6 \u03b8 \u03b8= + + 0 sin ,s \u03c6= 01 1 sin( ),s \u03c6 \u03b8= + and 012 1 2 sin( )",
            " While our inverse kinematic approach has a similarity with [4], there are difference as well in some respects such as: (i) due to the virtual joint based kinematic model, the forward kinematic Jacobian in our approach can be easily determined while the kinematic Jacobian in [4] was complicated to be obtained, and (ii) we do not carry the steering wheel velocity in the inverse kinematics but select it by an intuitive law, whereas, in [4], the steering law depended on the secondary optimization index selected for the null motion, not being selected independently. Virtual Joint Method for Kinematic Modeling of Wheeled Mobile Manipulators 1065 To verify the validity of the proposed approach, we present two simulation examples: (i) trajectory tracking tasks by car-like WMMs, and (ii) a collaborative object handling by three differential-drive WMMs. 5.1. Trajectory tracking by a car-like wheeled mobile manipulator In this simulation, we first conduct 2D trajectory tracking tasks using a planar car-like WMM with a two-link robot arm shown previously in Fig. 5; specifically, in this case, we consider the arm is placed at the front side in the central platform axis such that d = D and h = 0. Numerical values for the wheel-base and link lengths are chosen as: 1 2 1D L L= = = m. The virtual joint based Jacobian of this WMM becomes the same as the one in the case example in Section 3.3 for d = D and h = 0. We consider the tasks of line trajectory tracking. The line trajectories are created by the fifth-order polynomials in terms of the time parameter. These tasks are kinematically controlled by the inverse kinematic scheme in (20) and (21), where the manipulability measure is used as the secondary objective that defines the null motion of the mobile manipulator; please refer to Appendix A for the closed-form expression of the square of manipulability of the considered WMM, which is shown to be very concise thanks to the virtual joint based modeling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001838_978-3-319-19303-8_18-Figure18.13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001838_978-3-319-19303-8_18-Figure18.13-1.png",
        "caption": "Fig. 18.13 Concept of the H2S MEMS variable resistance sensor",
        "texts": [
            " The H2S sensors require high sensitivity to fairly low levels of the gas and must also be able to discriminate H2S from other gases that may be present and not give spurious readings affected by such other gases. Just as in many breathalizers, to operate, the H2S sensor\u2019s surface has to be heated well above ambient temperature to about 300 C. This puts a significant strain on the sensing module power supply and also prolongs a response time of the sensor. To reduce both, a modern MEMS technology may be employed. Figure 18.13 illustrates a concept of the MEMS H2S sensor, where a silicon structure is formed with a substrate supporting a thin membrane having thickness on the order of 1 \u03bcm. Such a membrane has low thermal capacity and thus can be warmed up to high temperature in a short time by a relatively low power. Two inter-digitized (alternating) electrodes are formed on the membrane\u2019s upper surface. A selective coating is deposited by a sputtering and oxidizing technique on the top of the electrodes. The coating has a finite resistance that varies in relation with concentration of the H2S molecules in the gas sample",
            " Other materials have been used to broaden the range of detectable chemicals, sol\u2013gel chemicapacitors, for example, can detect carbon dioxide [43]\u2014although such materials often have to be heated to achieve optimal performance. More recently, polymers have been used to make low power sensors for volatile organic compounds (VOCs) [44]. Chemicapacitors can be constructed using conventional thin film techniques, where conductive electrodes are arranged in either a parallel or interdigital layout, similar to Fig. 18.13. Typically, interdigitized electrodes consist of a single layer of metal deposited on a substrate to form two meshed combs. The polymer or other materials are deposited on top of the combs. Parallel-plate sensors [45] typically consist of a layer of metal deposited on a substrate, followed by a layer of insulator and finally a second, porous layer of metal on top of the insulator. One example of a chemicapacitor is a MEMS-sensor based on micromachined capacitors [46] has been developed and commercialized"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000854_iraniancee.2014.6999650-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000854_iraniancee.2014.6999650-Figure3-1.png",
        "caption": "Figure 3: 2-D finite element model of a BLDC motors with 14 poles and 12",
        "texts": [
            " By simplifying equation 4 is: ,0 5 Where lcm(x,y) is the least common multiple of its arguments, and q is a nonzero integer that results integer n. From the equation 5 Cogging torque harmonics are zero for witch is zero. Based on the above equation the fundamental harmonic of Cogging torque is , multiple of electrical frequency of motor. For the case of 14 poles and 12 slots the fundamental harmonics of Cogging torque is 6f. 978-1-4799-4409-5/14/$31.00 \u00a92014 IEEE 826 III. EFFECT OF POLE EMBRACE ON COGGING TORQUE The cross section of a BLDC motor with out-runner structure is shown in fig3. The motor has considered in this study has 12 slots and 14 poles. The approach used in this paper involved the use of a two-dimensional finite element package (MAXWELL 14). The field problem was made two dimensional by neglecting any end effects. This assumption might affect the magnitude of the torque in a short rotor, but is not expected to have significant influence on its variation with angle of rotation. [6] Figure1: Cross sectional a 14 poles motor with out-runner structure The specifications of the motor were considered here are shown in table 1. Pole embrace is the ratio of the pole width to rotor circumference. The cogging torque in figure 2 was obtained for motor which have 0.74 pole embrace. Figure2: cogging torque for motor with o.74 pole embrace Table 2 shows the cogging torque for various pole embrace; in motors have 14 poles and 12 slots. IV. UNBALANCED MAGNETIC FORCES The mechanism of unbalanced magnetic forces arising in BLDC motors is explained in [7] and [8]. Figure 3 shows the 2-D (FE) model of an-14 poles and 12-slots BLDC motor used for calculation of magnetic forces. slots Since the BLDC motor considered here has 14 poles, each poles of permanent magnet spans 25.7143 degree. Unbalanced magnetic force exists as long as there is eccentricity between the rotor and stator, because a portion of the stator is closer to the permanent magnet of the rotor, thus generating a net attraction force acting on the rotor [2]. For the eccentric rotor harmonics of the unbalanced magnetic forces are 1 where is the number of slots in the stator, is a positive integer and X denotes the rotational frequency of the motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure7-1.png",
        "caption": "Fig. 7 Placing the movement on ellipse path onto circle path",
        "texts": [
            " That is why the angular velocity fluctuation and pitch changing do not occur, namely: wP \u00bc wg \u00bc const: \u00f018\u00de As a consequence of this: rPe \u00bc rm \u00f019\u00de Figure 6 shows if we intersect the worm with an axial plane which is at the half of the lathe fork, then the path curve of the driving pin is an ellipse and the section of the driving pin is an ellipse curve as well. Thus, the ellipse cross section is moving on an ellipse path. Based on the CDE rectangular triangle, the major axis of ellipse is (Fig. 6): dcsen \u00bc dcs cos d1 \u00f020\u00de The minor axis of ellipse is: dcsek \u00bc dcs \u00f021\u00de In order to eliminate the angular velocity fluctuation, the moving ellipse has to be placed onto a radial circle path rm (Fig. 7). The displacement distance on the major radius of the ellipse is (Fig. 7): rhn \u00bc rne rm \u00f022\u00de The calculated displacement distance for the arbitrary ellipse P point is (Fig. 7): rhP \u00bc rPe rm \u00f023\u00de The polar equation of the ellipse section of the driving pin is (Fig. 7): xcs \u00bc rcs sin 8 1 ycs \u00bc rcsen cos 8 1 \u00f024\u00de Using the given equations above, we have prepared our own computer program (Figs. 8, 9, 10, 11, and 12). As the program knows the following pieces of information: the geometrical data of the worm, the driving pin diameter, the distance between the lathe fork and the worm shaft neck, the worm half cone angle, and the distance between the center line of the driving pin and the rotational axis of spindle (Fig. 3), then it calculates and represents the angular veloc- ity fluctuation \u03c9p as a function of spindle angular rotation \u03c6p (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001843_msf.836-837.177-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001843_msf.836-837.177-Figure2-1.png",
        "caption": "Fig. 2. (A) Machining strategy: full immersion slotting with constant depth of cut equal to 30 \u00b5m; tool geometry: (B) top view, (C) detail of the cutting edge radius, (D) tool dimensions",
        "texts": [
            " The experiments were carried out on an ultra-precision, Kugler\u2122 Micromaster 5-axis micro-milling machine, equipped with a granite structure, in an airconditioned room, to ensure constant temperature conditions. The machine is also equipped with an eddy current sensor that permits on-line control and compensation of the possible expansion of the air bearing spindle, which can reach speeds up to 180 000 RPM. The chosen machining strategy was the full immersion slotting with a slot width of 0.3 mm, length 25 mm and constant depth of the cut equal to 30 \u00b5m. The scheme of the slot geometry is shown in Fig. 2. All the experiments were conducted under dry lubrication conditions, which is particularly interesting for biomedical applications, to favour the component cleanliness. A full factorial design of experiments was used with two factors, namely the cutting speed and feet per tooth with two and four levels, respectively (see Table 1). Each experiment was repeated twice. diameter of 0.3 mm and cutting edge radius of about 1.1 \u00b5m, which leads to the critical feed per tooth of 0.33\u20130.44 \u00b5m/tooth (see Fig. 2). Each tool was inspected before machining to avoid any possible defects due to its fabrication. Materials Science Forum Vols. 836-837 179 Surface topography and roughness. Fig. 3 shows images of the slots bottom acquired by using a FEI\u2122 Quanta 400 SEM and a Sensofar\u2122 PL\u00b5 Neox confocal optical profiler. Regardless of the cutting speed and material microstructure, we can appreciate the same trend in the surface topography as a function of the feed per tooth. For the lowest value of the feed per tooth, the surface quality is very poor, due to the ploughing mechanism that prevails over shearing, whereas, when increasing the feed per tooth, the feed marks become more regular"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001800_s11432-015-5505-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001800_s11432-015-5505-5-Figure1-1.png",
        "caption": "Figure 1 (Color online) The proposed variable thrust direction mechanism.",
        "texts": [
            " In this paper, we develop a novel VTD mechanism to the conventional propeller engine UAV, which allows redirecting parts portion of the thrust from the propeller engine to other directions rather than normal axial direction. A combination flight controller for the VTD enhanced UAV is then proposed to coordinate the VTD controlled forces and aerodynamic surfaces forces. Finally, we verify the performance of the proposed VTD technique, in terms of the maneuverability, via both simulations and real flight test. The proposed variable thrust direction mechanism is shown in Figure 1, a conventional propeller engine is mounted on a two dimensional rotate disk, which is driven by two servo actuators. By combining the linear motions of the actuator, both the azimuth and the altitude angle of the disk with respect to the fuselage can be controlled, and thus changing the thrust direction of the propeller. As shown in Figure 2, the control schedule was employed that both the conventional control loop for aerodynamic surfaces and a VTD controller were used. Within conventional longitude control loop the error between desired height and actual height was used to drive the elevators and keep height constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001134_pvp2013-97863-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001134_pvp2013-97863-Figure2-1.png",
        "caption": "Figure 2: (a) Meshing of bolted flange and gasket, (b) Applied boundary conditions",
        "texts": [
            " Keeping in mind the rotational and reflective symmetry of the gasketed bolted flanged pipe joints, only one pipe, flange and half of the gasket is modeled. All of the flange and bolt dimensions and ratings are in accordance with ANSI B16.5 and gasket dimensions are as per ANSI B16.2 [25] Class 900#. ANSYS SOLID45 elements are used for both the flange and the bolt. Interface elements (INTER195) are used for the gasket. Contact elements, CONTA171 and CONTA174 are used to specify surface-to-surface contact pairs. The flange joint assembly with a suitable mesh of flange, bolt and gasket is shown in Fig. 2a. ANSYS software for finite element analysis is used [26]. Elasto-plastic material model is used for pipe, flange and bolt. Material properties are taken for flange and pipe as per [27] and are given in Table 1. linear loading and unloading curve using simplified model 1 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 06/19/2015 Terms of Use: http://asme.org/terms developed by Takaki et al [9] and used by Abid et al [21-24] and shown in Fig. 1. The flange and the gasket are free to move in the axial and radial direction. This provides flange rotation and the exact behaviour of stress variation in the flange, bolts and gasket. Symmetry conditions are applied to the gasket lower portion. An axial displacement is applied to the bolt bottom in downward direction to initiate contact and then to create the desired preload. Structural boundary conditions are shown in Fig. 2b. Bolt tightening of the 8 inch gasketed flange joint is performed in four pass incremental target stress for two different values used as per ASME PCC-1 guideline procedures [28] and using the following tightening sequence. T1 and T2 (T1=202 and T2=137MPa) are taken from an industry standard specification [29];  Sequence-1: Pattern style-1, 7, 4, 10, 2, 8, 5, 11, 3, 9, 6, 12 (for first three passes)  Sequence-2: Circular style-1, 2, 3, 4, 5, 6, 7, 8, 9, 10, 11, 12 (for the last/4th pass) The bolt preload scatter depends on the initial target torque applied to all the bolts in the joint assembly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002213_075045-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002213_075045-Figure1-1.png",
        "caption": "Figure 1. Pressure distribution in sealing contact areas of the piston rod seal.",
        "texts": [
            " For a static seal, the maximum contact compression stress of the seal should be larger than the working pressure, namely: ( )s p 1pmax where, p is oil medium pressure, \u03c3pmax is maximum contact compression stress. So, \u03c3p can be used as an important indicator to measure whether the static seal has failed or not. For a dynamic seal, because a complete work cycle should include outstroke and instroke, so the net leakage of rod seal is the leakage difference between the instroke and outstroke. A complete work cycle of piston rod seal which includes outstroke and instroke is shown in figure 1. The piston rod extends from the pressure cavity to the air side with a speed of uo and retracts in the opposite direction with a speed of ui. The cavity pressure is po , the height of the variable oil film is h(x), the variable pressure of liquid film is p(x), fluid viscosity between the gaps is \u03b7, the oil film thickness in outstroke and instroke are ho and hi, the oil film pressure at each maximum stress gradient of extending and instroke are wA and wE, respectively. For a stroke length of H, the net leakage of each circle is: ( ) ( ) ( ) p p h= - = \u22c5 - V dH h h dH u w u w 2 9 2 l o i o A i E From (2), it can be found that the leakage of the piston rod seal is determined by -u w u w ",
            " In the simulation model, the working pressure is given as 2 , 3 and 4MPa respectively, while 5% of the O ring is cut off to simulate wear of the seal. The contact strain distribution comparison between the normal ring and worn ring at 4 MPa is shown in figure 2. From figure 2, it can be found that when the seal is worn, the position of maximum contact strain inside the seal also changes. Because the material of the seal groove is 45 steel, its elastic modulus is much larger than that of the seal, the contact strain of the seal groove is much smaller than that of the seal, and so it cannot be shown in figure 2. From figure 1 in the above section, it can be found that no matter whether the cylinder is in instroke or outstroke, the contact pressure in the middle of the contact surface will be greater than both sides. So we choose the same point 1 in the middle of the contact surface on the seal groove, and compare the contact strain changes under different conditions. Results are listed in table 1. From table 1, it can be seen that the contact strain of the seal increases with the fluid pressure, but the contact strain of the normal seal is larger than that of the worn seal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003693_icem49940.2020.9270836-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003693_icem49940.2020.9270836-Figure1-1.png",
        "caption": "Fig. 1. SLM printing system: laser (a), scanning mirrors (b), protective glass (c), re-coater arm (d), build platform with the printed part (e), powder supply opening (f), powder coating (g)",
        "texts": [
            " SELECTIVE LASER MELTING SLM is a widely employed powder-bed based additive manufacturing process for rapid prototyping and the fabrication of highly specialized metal parts. The process involves the melting of thin metal powder layers with a focused laser beam. The layer thickness is usually within the range of 20-60 \u03bcm and the maximum laser powder in the range of 100 \u2013 1000 W. After melting each cross-sectional layer of the printed part, the build platform is lowered, additional powder is supplied on the platform, which then again is melted with the laser beam. The schematic of the printing system is presented on Fig. 1. The technology allows for the realization of three-dimensionally topology optimized parts with simplified logistics and smart material utilization, however its industrial scale application in the future would require considerably faster and more competitively priced SLM systems [11]. This research work has been supported by Estonian Ministry of Education and Research (Project PSG137). H. Tiismus, A. Kallaste, T. Vaimann and A. Rassolkin are with the Department of Electrical Power Engineering and Mechatronics, Tallinn University of Technology, Estonia (e-mail: hans"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.80-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.80-1.png",
        "caption": "Figure 4.80 Linear motion; roll feed drive",
        "texts": [
            " The unsupported weight of the material being moved must be included with the inertia of the rollers. 2. The spool of material from which the material is being removed may have to be accelerated by the roll feed. 3. The force which is used to load the two rollers together can cause an increase in the bearing torque above the normal free wheeling bearing torque. 4. Due to mounting tolerances and material compression, this force will not be perfectly normal to the center line of the roller bearing axis, and will create an additional load torque. See Figure 4.80. System Data: JM = motor inertia (g cm s2) JRM = motor roller inertia (g cm s2) JRP = pinch roller inertia (g cm s2) W = load weight (g) FL = load force (g) FF = friction force (g) TP = pressure torque (g cm) TB = bearing torque (g cm) DRM = motor roller diameter (cm) JSUPPLY = supply reel inertia (g cm s2) System Design 249 Material weight reflected to motor as an equivalent inertia: JW = ( W 980.6 ) ( DRM 2 )( 1 e ) Motion: Position: \u03b8M = S DRM/2 (rad) Velocity: \u03b8 \u2032 M = S\u2032 DRM/2 (rad s\u22121) Acc/Dec: \u03b8 \u2032\u2032 M = S\u2032\u2032 DRM/2 (rad s\u22122) At Motor: TR = (FL + FF ) ( DRM 2 )( 1 e ) + TP + TB (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003115_physrevfluids.5.053103-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003115_physrevfluids.5.053103-Figure5-1.png",
        "caption": "FIG. 5. Uniformly seeded tracers mixed by cilia in a wavy channel. (a) Initial seeding. (b) Tracers after ten cycles for Nw = 10 ( \u03c6 = 10\u03c0/16). (c) Mixing efficiency after ten cycles as a function of Nw .",
        "texts": [
            " The coefficient in the denominator is to scale the channel such that the fluid domain has the same area compared to the regular circular channels. The coefficient of the cosine term perturbs the radius of the outer boundary by about \u00b110%. In other words, the narrowest and the widest channel widths are about 1.5 and 2.5 unit length. The initial seeding and the tracer positions after ten beating cycles are shown in Figs. 5(a) and 5(b) with Nw = 10, which yields the best mixing results as shown in Fig. 5(c). When compared to Fig. 4, it is clear that not only the number of waves that yields the best mixing performance changes from \u22129 to 10, but also the overall mixing performance is negatively affected by the presence of the wavy channel, indicated by the mixing number ln(m/mo) increase from \u22123.4 to about \u22122.8 (in other words, m/mo increased from e\u22123.4 \u2248 0.03 to e\u22122.8 \u2248 0.06). The effect of the wall perturbation on mixing is even apparent to the eye: in Fig. 5(b), at each of the humps on the outer wall, a shear region could be observed which does not exist in the case of the regular Taylor-Couette geometry. In this section we study the full cilia-channel-particle problem and compare the results with passive tracers, in an effort to showcase the effects of the particle size in such problems. We start by uniformly seeding 20 circular particles of radius rp inside the channel and trace their centroids within one ciliary beating cycle. With small particle size, as shown in the top row of Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001405_j.mechmachtheory.2014.11.017-Figure8-1.png",
        "caption": "Fig. 8. Relative orientation of the five bodies in the tangent plane. n points out of the plane of the paper.",
        "texts": [
            " Tmnpm and Tfnpf are the instantaneous Darboux frames [21] on SFm and SFf respectively at the point of contact. A virtual five-link mechanism consisting of the f-surface (body-1 in [15]), them-surface (body-2 in [15]), and three virtual bodies: Tnp based virtual body (body-3 in [15]), Tfnpf based virtual body (body-4 in [15]), Tmnpm based virtual body (body-5 in [15]), is used to derive the contact kinematics between the two surfaces. The relative orientation of all the frames in the tangent plane is shown in Fig. 8.\u03b1m and\u03b1f are the spin anglesmade byT withTm andT f respectively in the sense ofn. The instantaneous relative velocity between the surfaces at the point of contact C is directed alongT . \u03b2 is the angle made by T with positive X-axis. Let \u1e61m and \u1e61f be the speeds along the contact traces on the two surfaces. S \u00bc S \u00fe \u03f5So and S \u00bc S \u00fe \u03f5So are the twist and differential twist dual vectors respectively. The contact kinematics equations given by Eq. (20) in [15], but written using the current notation are: 0 \u00bc s mt m\u2212s f t f \u00fe S T \u00f011\u00de 0 \u00bc s m\u03b3 m\u2212s f\u03b3 f \u00fe S n \u00f012\u00de 0 \u00bc s mk m\u2212s f k f \u00fe S p \u00f013\u00de 0 \u00bc s mc\u03b1m \u2212s f c\u03b1 f \u00fe vc \u00f014\u00de 0 \u00bc s ms\u03b1m \u2212s f s\u03b1 f \u00f015\u00de 0 \u00bc S o n \u00fe s f c\u03b1 f S p\u2212s f s\u03b1 f S T\u2212s f k f vc \u00f016\u00de 0 \u00bc S o p\u2212s f c\u03b1 f S n \u00fe s f\u03b3 f vc \u00f017\u00de where tf\u204e, \u03b3f \u204e, kf\u204e, tm\u204e, \u03b3m \u204e, and km\u204e are given by: t f \u00bc t f c\u03b1 f \u2212kf s\u03b1 f \u00f018\u00de \u03b3 f \u00bc \u03b3 f \u00fe \u03b10 f \u00f019\u00de k f \u00bc t f s\u03b1 f \u00fe kf c\u03b1 f \u00f020\u00de t m \u00bc tmc\u03b1m \u2212kms\u03b1m \u00f021\u00de \u03b3 m \u00bc \u03b3m \u00fe \u03b10 m \u00f022\u00de k m \u00bc tms\u03b1m \u00fe kmc\u03b1m \u00f023\u00de where tm, km, and \u03b3m are the geodesic torsion, normal curvature, and geodesic curvature respectively of SFm at the point C in the direction of Tm "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001118_j.gene.2013.03.006-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001118_j.gene.2013.03.006-Figure1-1.png",
        "caption": "Fig. 1. Number of HPRT+ (IMPPase-positive) transformants per ml of \u201creaction mixture\u201d, assayed in HAT medium, following treatment of HPRT\u2212 (IMPPase-negative) D98/AH2 recipient cells with increasing concentrations of DNA isolated from HPRT+ (IMPPase-positive) D98S donor cells. (From: Szybalska EH, Szybalski W. 1962. Genetics of human cell lines. IV. DNA mediated heritable transformation of a biochemical trait. Proc Natl Acad Sci 48: 2026\u20132034. Copyright by the authors).",
        "texts": [
            " 3: 272) and were not able to form colonies on HAT medium (Szybalska EH, Szybalski W. 1962. Genetics of human cell line. IV. DNA-mediated heritable transformation of a biochemical trait. Proc Natl Acad Sci. 48: 2026\u20132034). These cells, and especially the D98/ AH-2 line, which had never produced any HPRT+ revertants that grew on the HAT medium, were exposed to DNA purified either by the phenol method or by CsCl density gradient centrifugation. When freshly Ca phosphate-precipitated DNA was used, colonies were formed over a 100\u20131000-fold DNA concentration range (see Fig. 1). DNA isolated from transformants had a similar transforming (transducing) activity (T.A.) as that extracted from the wild-type Detroit-98 cells (see Fig. 2). DNA isolated from HPRT\u2212 (IMPPase-negative) D98/ AH2 had no transducing activity. The transforming activity (T.A.) was insensitive to RNases, but sensitive to DNases. In CsCl gradient, T.A. banded at the same density as human DNA (1.7 g/ml). Similarly, in Cs2SO4 gradient, T.A. banded at the same density as human DNA (1.4 g/ml). About two years after our discoveries of 1962, a neurological \u201csyndrome\u201d caused by the loss of the HPRT enzyme was discovered in children by Lesch and Nyhan (Lesch M, Nyhan WL"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001548_iciea.2015.7334299-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001548_iciea.2015.7334299-Figure9-1.png",
        "caption": "Fig. 9. Magnetic field lines",
        "texts": [
            " Equation(11) shows eddy current loss is proportional to the square of alternating frequency and the square of maximum flux density , is inverse proportional to the resistivity , and related to the structural parameters of PM. B. Simulation model A Two-dimensional model of FSCW-PMSM with 24-slot 16-pole is built by using FEM. On the basis of cycle symmetry, the whole model is divided into 1/8, and it is shown in Fig.6. Considering different area of the motor and the influence of skin effect, the grid subdivision shown in Fig.7. Fig.8 and Fig.9 represent the distribution of the flux density and the magnetic field lines on the motor model. The PM eddy current loss under rated-load condition is shown in Fig.10. It fluctuates periodically, because it changes with the change of the stator coil position. 1248 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) C. Eddy current loss of main harmonics The fundamental harmonic of air gap field has the same speed as rotor, so it does not produce eddy current on PMs. But the speed of other harmonics is different from rotor, the speed of times harmonic is of rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003708_5.0015728-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003708_5.0015728-Figure1-1.png",
        "caption": "FIGURE 1. Equivalent Stress Simulation Results",
        "texts": [],
        "surrounding_texts": [
            "The first step in this research is to do an image design using Autodesk Inventor software. Then, do a simulation with Ansys software. The type of analysis used in this software is static structural analysis and explicit dynamic analysis, where input material properties are obtained from AISI 4315 mechanical properties. In the final stage, the expected results of the finite element method simulation are equivalent stress, equivalent elastic strain, total deformation, stress intensity factor and J-integral."
        ]
    },
    {
        "image_filename": "designv11_30_0003540_s1064230720050081-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003540_s1064230720050081-Figure2-1.png",
        "caption": "Fig. 2. Two-link mechanism: simplified (auxiliary) model of person on swing.",
        "texts": [
            "11) have the dimension of the moment of inertia (see (2.2)). In addition to the control moments L and M, Eqs. (2.7) and (2.8) also take into account the moments LS and MS (see formulas (1.2)). In order to build the control (in the form of feedback) in the three-link swing model, first, we consider an auxiliary (simplified) swing model that contains only two links. By exploring this simpler model, it is possible to build the quasi-optimal control, which is then used to control the more complex three-link model shown in Fig. 1. Figure 2 shows a relatively simple model of a swing with a person on it. As mentioned above, such a model has been considered in many works. The fixed hinge O is the suspension point of the swing OC, which is considered to be weightless, and l is its length. As in the three-link model, we denote by \u03d5 the angle of the swing\u2019s deviation from the vertical. A person is modeled by the absolutely solid homogeneous rod HF, whose center of mass C is hinged to the free end of the swing. We denote by \u03b8 the angle between the continuation of the swing OC and the rod HF",
            "h h s s s b b bb m gOC b m gOK b m gKC b m gOP b m gPC P bI O hI K sI \u03b4 = \u03b4\u03b1 + \u03b4\u03b2,W L M ( ) ( ) \u03b1 \u03b2 \u03d5 + \u03b1 \u03b1 + \u03b2 \u03b2 \u2212 \u03d5\u03b1 \u03b1 \u2212 \u03b3 \u2212 \u03b1 \u03b1 \u2212 \u03b3 + \u03d5\u03b2 \u03b2 + \u03b4 + \u03b2 \u03b2 + \u03b4 = \u2212 \u03d5 \u2212 \u03d5 + \u03b3 \u2212 \u03d5 + \u03b1 \u2212 \u03d5 \u2212 \u03b4 + \u03d5 + \u03b2 \u2212 \u03bc\u03d5 2 1 2 3 12 12 2 13 13 1 2 3 4 5 ( , ) ( ) 2 sin( ) sin( ) 2 sin( ) sin sin sin( ) sin( ) sin( ) sin( ) , j j j a a \u0430 a b b b b b ( ) ( ) ( )\u03b1 \u03d5 + \u03b1 + \u03d5 \u03b1 \u2212 \u03b3 = \u2212 \u03d5 + \u03b1 + + 2 2 22 12 3sin sin ,Sj a a b L L ( ) ( ) ( )\u03b2 \u03d5 + \u03b2 \u2212 \u03d5 \u03b2 + \u03b4 = \u03d5 + \u03b2 + + 2 3 33 13 5sin sin .Sj \u0430 a b M M ( ) ( ) ( )\u03b1 \u03b2 = + + + \u03b1 \u2212 \u03b3 \u2212 \u03b2 + \u03b41 11 22 33 12 13, 2 cos 2 cos ,j a a \u0430 a \u0430 ( ) ( )\u03b1 = + \u03b1 \u2212 \u03b32 22 12 cos ,j a a ( ) ( )\u03b2 = \u2212 \u03b2 + \u03b43 33 13 cos .j a a = 2/12\u0421J mr JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 59 No. 5 2020 relative to swing OC (Fig. 2). Moment Q, as well as the moments L and M (see (1.1)), is considered to be limited in magnitude: , where . The equations of motion of the two-link mechanism described above are presented as follows: (3.1) Here, is the moment of inertia of the two-link mechanism relative to the suspension point O, while describes, as in Eq. (2.6), the moment of viscous friction forces at point O. In hinge C, in addition to the control moment Q, moment QS is also applied; it is designed to prevent violation of the conditions (3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002033_icelmach.2014.6960180-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002033_icelmach.2014.6960180-Figure8-1.png",
        "caption": "Fig. 8. Time-harmonic FEA simulation result with rotor conductivity set equal to \u03c3=5 MS/m.",
        "texts": [
            " However, along the periphery of rotor slots and around the outer rotor circumference, the local conductivity should be set to a much higher value\u2212 approaching MS/m as per hypothesis a) \u2013 due to the local short circuits between adjacent silicon sheets caused by manufacturing defects (punching and cutting). As a consequence, in presence of the mentioned uncertainty, FEA simulations have been repeated according to the two \u201cextreme-case\u201d scenarios a) and b), i.e. assigning the rotor core a uniform conductivity respectively equal to =5 MS/m and =0.021 MS/m. FEA time-harmonic simulation output proved to be highly sensitive to the rotor conductivity value being set in terms of rotor flux distribution, as it can be clearly seen from Fig. 8 and Fig. 9, which refer to the same value of frequency and current. In fact, for the higher rotor conductivity it is clear that a large amount of reaction current arises in the rotor core (as if it were a solid-steel one) causing a sort of \u201cmagnetic shielding\u201d effect which rejects the flux lines towards the outer rotor periphery according to the well-known \u201cskin effect\u201d (Fig. 8). However, it will be seen in the following that the sensitivity of rotor parameters to the rotor axial conductivity values is not so high as could be expected looking at the difference flux distribution (Figs. 8-9) since, regardless of the conductivity value being set, the vast majority of the rotor current reaction comes from the cage and not from the eddy current in the core. For the computation of this parameter, reference is made to the simplified equivalent circuit shown in Fig. 10 where Isc is the short-circuit current and Vsc\u2019 is the stator emf due to the total air-gap flux except for stator end-coil leakage flux contribution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003336_02640414.2020.1798716-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003336_02640414.2020.1798716-Figure1-1.png",
        "caption": "Figure 1. Static shooting accuracy (SSA) assessment consisted of trials from the single position at central position (3), while Dynamic shooting accuracy (DSA) assessment consisted of trials from five positions: 1) full left \u2013 left corner, 2) half left \u2013 position between left corner and central position, 3) central position \u2013 in front of the basketball table, 4) half right \u2013 position between left corner and central position and 5) full right \u2013 right corner; while the start was from the left-to-right corner.",
        "texts": [
            " The subjects were asked to follow their normal diet and to refrain from any form of intense physical activity for the 48 hours, as well as to fast for 2 hours before each testing session. The subjects wore standard basketball clothing and shoes on testing. Assessments of shooting accuracy performance consisted of static shooting accuracy (i.e. SSA; free-throws or static shots) and dynamic shooting accuracy (i.e. DSA; jump-shots or dynamic shots). Assessment of SSA consisted of block with 10 static trials (i.e., free-throws or static shots from the central position in front of basketball table; see Figure 1), while DSA assessment included block of 15 dynamic trials (i.e. jump-shots or dynamic shots), that were split into 3 consecutive trials from each of 5 standard positions (for details see, Figure 1) performed with time limitation of 50 seconds (Karaleji\u0107 & Jakovljevi\u0107, 2009). Each shooting accuracy testing sessions were preceded by a standard 10-min warming up and 5-min active stretching procedure and followed by detailed explanations and qualified demonstrations of each test. After 15 minutes warm up using standardized exercises and a block of 10 practice trials was performed as a special warm up for SSA and DSA testing. The participants were instructed to perform either SSA or DSA trials as accurately as possible regardless of the experimental conditions",
            " Note that experimental groups used only one ball size depending of groups they belonged (i.e. EGchildren ball, EGadult ball). The shooting was performed at children basket height (i.e. 2.6 m) with appropriate balls and included SSA (30 static trials in total) and DSA (90 dynamic trials in total) practice. Specifically, SSA practice consisted of 3 block of 10 trials performed at single shooting distance of 2.7 m, while DSA practice involved 3 block of 30 trials, whereas 1 block consist of 2 dynamic trials performed on each of 5 positions (for details see, Figure 1) and at 3 shooting distances: 2.0 m, 2.7 m and 3.4 m. During each training session, SSA and DSA practice was changed after one block of trials, alternately. The rest periods among consecutive trials were 3 to 5 seconds; while between 2 consecutive positions and 2 shooting distances were 10 and 30 seconds, respectively. The pause between different types of practice was approximately 1 minute. To assess shooting performance of SSA and DSA the amount of error within the block of trials was recorded using dichotic outcomes (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001355_gt2013-95816-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001355_gt2013-95816-Figure2-1.png",
        "caption": "FIGURE 2: TOUCHDOWN BEARING MODULE",
        "texts": [
            " Inside the ball bearings, contacts between balls and inner respectively outer bearing races, as well as ball-ball contacts are assured. The link between bearing outer rings and the stator is realized by a compliant damping element, in reality embodied by a special ribbon with non-linear behavior. It is radially compressible over a certain distance, but once completely crushed the stiffness increases. In the model this is implemented by a second parallel spring-damper element responding after a certain clearance (c.f. Figure 2). The simulation runs are generally organized in three steps: - At still stand, bearing preloads are set; all elements need to get placed in their initial position - The levitated rotor is speed up; unbalance is adjusted - The actual drop is initiated by disabling the links holding the rotor in levitation The modeling technique was applied on three different real machines. The validation was done thanks to rotor drop test campaigns. Test rig 1: \u201c4096-Rig\u201d The 4096-Rig is a small horizontal turbo-compressor for a train air conditioning application"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002254_s0081543816060171-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002254_s0081543816060171-Figure11-1.png",
        "caption": "Fig. 11. Phase portrait (section formed by the intersection with the plane g = \u03c0/2) in Andoyer\u2013 Deprit variables (L/G, l) for the Hess case under the conditions I = diag(1, 0.625, 0.375), r = (3, 0, 4), and \u03bc = 1.995 with constant integrals h = 50 and f = 5. Two stochastic layers divided by the double Hess separatrix are well visible: points from one layer do not penetrate into the other. In (b) a meandering torus arising under these conditions can also be seen (see Fig. 12).",
        "texts": [
            " 294 2016 Generally speaking, the dynamics on this invariant Hess manifold differs from the usual quasiperiodic motion, which arises when the conditions of the Liouville\u2013Arnold theorem are satisfied. The Hess case cannot in general be integrated by quadratures on (A.3) but nevertheless can be analyzed qualitatively. Remark A.1. In this appendix, we construct phase portraits by using the Andoyer\u2013Deprit variables (see [13] for details). For certain values of the energy and area integrals, the Hess relation can define a pair of double separatrices on the phase portrait (see Fig. 11) that separate two chaotic zones (which show that there exists no general integral under the Hess conditions). It is interesting to note that in the phase space a meandering torus arises for the Hess case (see Fig. 12). Such an effect is due to the loss of twisting and is encountered in Hill\u2019s celestial mechanics problem [46, 47] and in the planar restricted three-body problem [21]. Proposition A.1. The Hess case in the Suslov problem (considered in Subsection 2.4) is equivalent to the Hess case in the Euler\u2013Poisson equations",
            " 294 2016 where M1 = K sin l, M2 = K cos l, and \u03bc\u0303 = \u2212\u03bc. Write the Hamiltonian H = 1 2 ( M2 1 + 2 3 M2 2 + 1 2 M2 3 ) + 1\u221a 3 \u03b31 + 1\u221a 6 \u03b33. (A.16) There are two qualitatively different cases (this result was first obtained in the book [13]). Case h > \u03bc\u0303. The center of mass rotates in the principal circle (since \u03c8 = const). In this case the middle axis moves along the entire loxodrome, and on the phase portrait (Figs. 13e and 13f), which also contains chaotic trajectories, the Hess solution separates two \u201cimmiscible\u201d stochastic layers (see also Fig. 11). The actual Hess solution in this case is not implementable: due to instability the trajectory \u201cfalls down\u201d into one of these layers. As h \u2192 \u221e (or \u03bc\u0303 \u2192 0), everything reduces to the standard Euler case and the Hess solution tends to the separatrix of permanent rotation about the middle axis [32]. PROCEEDINGS OF THE STEKLOV INSTITUTE OF MATHEMATICS Vol. 294 2016 Case h < \u03bc\u0303. The center of mass executes flat oscillations according to the law of a physical pendulum, and the middle axis moves according to (A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003606_01691864.2020.1835532-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003606_01691864.2020.1835532-Figure2-1.png",
        "caption": "Figure 2. Basic ideas for base kinematics estimation. (a)Integration of the IMU output. (b)Leg kinematics. Under the assumption that the supporting foot is fixed, base kinematics is computed from the supporting foot.",
        "texts": [
            " However, they put two assumptions: one is that the CoM height is constant, and the other is that the moment around the CoM is negligible. Carpentier et al. [24] eliminated the former assumption by introducing the projection on the central axis of contact wrench. By using the central axis instead of the ZMP, their CF can estimate three-dimensional CoM. However, the latter assumption still remained. For this problem, Bailly et al. [33] proposed a simultaneous estimation of the CoM and the moment around the CoM. They showed that the moment estimation can improve the accuracy of the CoM estimation. This idea, shown in Figure 2(a), is well-known in the field of themobile robot. The IMUoutputs the linear acceleration and the angular velocity in the sensor frame. Suppose that we install the IMU on the base link, the base kinematics p0 and R0 are related with the IMUmeasurement, namely, p\u03080 = R0 0a0 \u2212 g, (30) R\u03070 = R0 [ 0\u03c90\u00d7 ] , (31) where 0a0 \u2208 R 3 and 0\u03c90 \u2208 R 3 are the linear acceleration and the angular velocity of 0 with respect to 0, respectively. [\u2217\u00d7] \u2208 R 3\u00d73 is the matrix which means the cross product by the three-dimensional vector \u2217",
            " The IMU measurements are often regarded to have the bias and the noise, namely, 0a\u03030 = 0a0 + 0ba + 0wa, (32) 0\u03c9\u03030 = 0\u03c90 + 0b\u03c9 + 0w\u03c9, (33) 0b\u0307a = 0wba, (34) 0b\u0307\u03c9 = 0wb\u03c9, (35) where 0a\u03030 and 0\u03c9\u03030 denote the measurement of 0a0 and 0\u03c90, respectively. 0ba \u2208 R 3 and 0b\u03c9 \u2208 R 3 are the bias of the acceleration and the angular velocity, respectively. 0wa, 0w\u03c9, 0wba, and 0wb\u03c9 \u2208 R 3 are the white Gaussian noise. Although p0 and R0 can be obtained by integrating (30) and (31), the error due to the bias and noise is accumulated. This idea is inherent to the legged robot, and p0 and R0 are computed by the kinematics from the support foot to the base-link frame 0, as shown in Figure 2(b). This kinematics is written as pS = p0 + R0 0pS, (36) RS = R0 0RS, (37) where pS \u2208 R 3 and RS \u2208 SO(3) are the position and attitude of the support foot with respect to , respectively. 0pS \u2208 R 3 and 0RS \u2208 SO(3) mean the relative position and attitude between the support foot and 0, respectively. Those relative value can be calculated when the measured joint displacements and the link property are available. If pS and RS are known, p0 and R0 can be obtained as p0 = pS \u2212 R0 0pS, (38) R0 = RS 0RT S , (39) Although pS andRS are fixed during the supporting phase in many conventional cases [37,46], this assumption is not always satisfied"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002729_tmag.2020.2966572-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002729_tmag.2020.2966572-Figure14-1.png",
        "caption": "Fig. 14. Data acquisition system: (a) the XY-axis motorized linear stages and the programmable Gaussmeter; (b) an illustration of the 122\u00d756 data matrix above the magnet.",
        "texts": [
            " A magnet holder is designed to secure the magnet samples and the Hall-effect sensor. When the holder is inserted in the test fixture, the magnet sample is placed in the center of the airgap. The easy axis (north polarization) of the magnet is opposed to the demagnetizing field generated by the test fixture 0018-9464 (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. as shown in Fig. 14(b). During the test, the Hall-effect sensor is fixed in the notch of the black sensor holder as shown in Fig. 2(b). The active area of the sensor (black dot marked in Fig. 2(b)) directly touches the center of the north pole of the magnet and a positive value of flux density can be measured. As a transverse Hall-effect sensor, it only records the flux density vector components perpendicular to the sensor, i.e., components along the easy axis of the magnet as shown the red dashed arrow in Fig. 2(b)",
            "html for more information. pyramid poles. Two 0.2-mm airgaps are left between the magnet and the two steel blocks. Each airgap is filled by two pieces of paper so that the magnet is not damaged by the steel. The 0.4-mm total airgap is much thinner than the 7.8-mm airgap in section \u2163. Therefore, the current required to demagnetize the magnet is much lower than that in section \u2163. A programmable Gaussmeter is used to sweep the flux density on the surface of the magnet after demagnetization as shown in Fig. 14 (a). The distance between the magnet surface and the active area of the Gaussmeter is 0.5 mm. The motion of the Gaussmeter is controlled by an XY-axis motorized linear stage. The step size of the linear stages is 0.2 mm in both directions. Each group of sweeping data is a 122\u00d756 matrix. These data are the components of flux density vectors along the easy axis as shown in Fig. 14(b). B. Test Results of Time-dependent Demagnetization Induced by Pulsed Demagnetizing Fields Two groups of tests were conducted to understand the timedependent demagnetization induced by pulsed demagnetizing field. In test one, pulsed fields generated by multiple levels of current are applied to the magnet samples. Test one follows the procedure below. 1. Assemble the test fixture. Power on the FPGA, the IGBT module and the gate drive. 2. Program the FPGA. Make sure the duty ratio in period one is 10%, and the duty ratio in period two is the designated value",
            " It can be estimated that the flux linkage of the magnet will cross the simulation result (red dashed line) at the 180th unit time. The flux linkage will continue decreasing if more demagnetizing pulses are applied. C. Spatial Flux Density Distribution of Magnets after Pulsed Contact Demagnetization In Section V.B, the flux linkage data only tell the demagnetization level of the magnet after pulsed field. However, they do not contain detailed flux density information over the entire magnet surface. The flux density acquisition system shown in Fig. 14 can measure the flux density of numerous points on the magnet surface. The spatial flux density data after pulsed demagnetization are visualized in this section. A group of magnet contours of sample 1 NdFeB magnet after multi-level non-contact (7.8-mm airgap) constant field demagnetization are shown in Fig. 17. As can be seen, with the increase of demagnetization level, the magnet center gets deeper demagnetization because the magnet center suffers a stronger demagnetizing field than the edge. This is one of the disadvantages of demagnetization tests with large airgap"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000918_icems.2013.6754383-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000918_icems.2013.6754383-Figure5-1.png",
        "caption": "Fig. 5. Leakage flux at rib in motor proposed.",
        "texts": [
            " 978-1-4799-1447-0/13/$31.00 \u00a92013 IEEE A. Torque Characteristic Fig. 3 and 4 show the relation between the average torque and the phase angle \u03b2 of the motor proposed and Type A, respectively. The maximum total torque is obtained when \u03b2 is 50 degrees because the reluctance torque is high as shown in these figures. The maximum torque of motor proposed is 56.1 N\u30fbm and that of Type A is 57.0 N\u30fbm. The structure of ribs in the motor proposed decreases the average torque because of the leakage flux as shown in Fig. 5. B. Torque and Output for Speed Fig. 6 and 7 show the torque and output for the speed, respectively. The motors are controlled by the field weakening method. The maximum output of the motor proposed shows 15.1 kW and that of Type A is 15.4 kW as shown in Fig. 7. The output of the motor proposed decreases due to the torque reduction. Fig. 8 shows the relation between the phase angle \u03b2 and the rotor speed. As the rotor speed becomes higher, the phase angle \u03b2 becomes larger to maintain a line voltage of 400 Vrms as shown in this figure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001167_s00170-014-6109-8-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001167_s00170-014-6109-8-Figure14-1.png",
        "caption": "Fig. 14 FEA model of cutting edge of surface broaching tool",
        "texts": [
            " In addition, enough tool edge strength of broaching tool is particularly important to improve tool life and machined surface quality. Furthermore, the structure strength of tool edge is another important constraint, which is also be determined by rake angle, relief angle, and RPT. The value of tool edge strength could be obtained by finite element methodology (FEM). In similar manner, the same CCF experiment design in the FEA simulation can also be applied to estimate the effects of RPT, rake angle, and relief angle on the strength of tool edge. As depicted in Fig. 14, the loads (Fx and Fy) measured by cutting experiments are applied on the nodes of cutting edge in FEAmodel. And the degree of freedom (DOF) of connected and fixed regions is set to zero (DOF=0). The regressionmodel of tool edge\u2019s structure strength S can also be obtained by RSM. The quadratic regression equation of maximum stress on the cutting edge is described as follows: S \u00bc \u22122:36\u03b32 \u2212 9:54\u03b12 \u00fe 22358\u03b42 \u2212 1:56\u03b3\u03b1 \u2212 11:0\u03b3\u03b4 \u2212 85:4\u03b1\u03b4 \u00fe 97:4\u03b3 \u00fe 123\u03b1 \u2212 2078\u03b4 \u2212 1053 \u00f09\u00de where \u03b4, \u03b3, and \u03b1 are RPT, rake angle, and relief angle, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003204_1045389x20935573-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003204_1045389x20935573-Figure8-1.png",
        "caption": "Figure 8. Magnetostriction phenomenon of Terfenol-D.",
        "texts": [
            " International Journal of Precision Engineering and Manufacturing 19: 495\u2013503. Weck M and Koch A (1993) Spindle bearing systems for high-speed applications in machine tools. CIRP Annals\u2014 Manufacturing Technology 42: 445\u2013448. Zhang T, Jiang C, Zhang H, et al. (2004) Giant magnetostrictive actuators for active vibration control. Smart Materials and Structures 13: 473\u2013477. Zverev IA, Eun IU, Hwang YK, et al. (2006) An elastic defor- mation model of high-speed spindle units. International Journal of Precision Engineering and Manufacturing 7: 39\u201346. Appendix 1 Figure 8 shows the magnetostriction phenomenon of Terfenol-D. The intensity of the magnetic field is expressed as equation (3) H =NI=l \u00f03\u00de where H is the magnetic field intensity, N is the number of coil turns, l is the device length, and I is the input current. Magnetostriction can be expressed as equation (4) l= cs +Hd \u00f04\u00de If s = 0, equation (4) can be expressed as equation (5) where l is magnetostriction, c is compliance, s is prestress, and d is the magnetic constant l=Hd \u00f05\u00de Also, magnetostriction is expressed as equation (6) l=Dl=l \u00f06\u00de Here, Dl is the extended device length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001087_0954409714530912-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001087_0954409714530912-Figure4-1.png",
        "caption": "Figure 4. A case of wheel/rail contact.",
        "texts": [
            " This calculation iteration continues until the coordinates y1 \u00bc 2:5227\u00fe 1 8 x 42:2840\u00f0 \u00de\u00bd , 60:05 x4 42:2840 220:8243\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2202 x \u00f0 32\u00de\u00bd 2 q , 42:28405 x4 20:1944 498:9749 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 5002 x \u00f0 32\u00de\u00bd 2 q , 20:19445 x49:5625 100:3593 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1002 \u00f0x 1:25\u00de2 q , 9:56255 x429:1531 17:775 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 142 \u00f0x 25:2466\u00de2 q , 29:15315 x438:4025 12:9872\u00fe x 38:4025 39:3543 38:4025 \u00f012:9872\u00fe 15:6023\u00de, 38:40255 x439:3543 9:4454\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 182 \u00f0x 56:2686\u00de2 q , 39:35435 x449:4629 15:0\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 122 \u00f0x 54:0\u00de2 q , 49:46295 x462:0 6:0557\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 242 \u00f0x 46:0\u00de2 q , 62:05 x470:0 8>>>>>>>>>>>>>>>>>>>>>>>>>>>>< >>>>>>>>>>>>>>>>>>>>>>>>>>>>: \u00f01\u00de y2 \u00bc 14:8408\u00fe h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 132 x l \u00f0 22:4108\u00de\u00bd 2 q , 35:39465 x4 25:3361 80:1223\u00fe h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 802 x l \u00f0 7:3333\u00de\u00bd 2 q , 25:33615 x4 10:0 300\u00fe h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3002 \u00f0x l \u00de2 q , 10:05 x410:0 80:1223\u00fe h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 802 \u00f0x l 7:3333\u00de2 q , 10:05 x425:3361 14:8408\u00fe h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 132 \u00f0x l 22:4108\u00de2 q , 25:33615 x435:3946 8>>>>>>>>>>>>< >>>>>>>>>>>>: \u00f02\u00de at RICE UNIV on May 18, 2015pif.sagepub.comDownloaded from of the contact point are found at a wheel/rail relative position. Figure 4 is a simplified schematic diagram that shows the point M where contact occurs between the wheel and rail. Parameter s is the horizontal distance between contact point M and the centerline of the rail\u2019s cross-section, and it is set to be positive when the contact point is on the right-hand side of the centerline of the rail\u2019s cross-section. The contact point and distance s at any wheel/rail contact point (different values of parameter l) can be obtained by repeating the calculation process. Figure 5 shows the flowchart of the calculation of the contact point under different conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure6-1.png",
        "caption": "Fig. 6. Principle of the suspension force Fb+.",
        "texts": [
            " The stator-laminated cores are inserted into a nonmagnetic material case from the axial direction, as shown in Fig. 5(a). In addition, small magnets are inserted to the holes of the stator-laminated cores. Subsequently, it is possible to assemble the stator easily by inserting the stator bulk cores with the windings to the predefined position of the stator-laminated cores, as shown in Fig. 5(b). Accordingly, the stator inner surface accuracy is high because of the unified stator laminated cores. Fig. 6 shows the principle of generating the suspension force Fb+ toward the positive direction in the b-axis. The x-axis and the y-axis are defined as the perpendicular coordinates fixed to the stator with reference to the winding position. In addition, the a-axis and the b-axis are converted from the x-axis and the y-axis using rotational conversion with 45\u00b0 toward the clockwise direction. The novel homopolar-type MB has four independent windings Nx+, Nx\u2212, Ny+, and Ny\u2212 to generate the suspension forces Fx+, Fx\u2212, Fy+, and Fy\u2212 toward the four different directions, respectively. The magnetic flux \u03c6x+ is generated by the current ix+ flowing in the Nx+ winding, as shown in Fig. 6. Then, the magnetic attractive force Fx+ toward the positive direction in the x-axis is generated in the gaps between the stator laminated cores and the rotor laminated cores. Similarly, Fy+ toward the positive direction in the y-axis is generated. Thus, the suspension force Fb+ can be generated toward the positive direction in the b-axis by a vector sum of Fx+ and Fy+. Based on the principle of generating the suspension force in Section III-B, the novel homopolar-type MB, shown in Fig. 4(a), is evaluated with the 3D-FEA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002808_0959651819899267-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002808_0959651819899267-Figure11-1.png",
        "caption": "Figure 11. Euler angle transformation function sub-model.",
        "texts": [
            " Euler angle is used in this work because the pitch angle (i.e. the Euler second rotation angle) would never be 90 . So, gimbal lock type of singularity does not arise. The transformation from bodyfixed angular velocities to Euler angles rates is given by16,21,49 _f _u _c 0 @ 1 A= 1 b1 b2 0 b3 b4 0 b5 b6 2 4 3 5 vx vy vz 0 @ 1 A \u00f013\u00de where b1 = tan u sinf, b2 = tan u cosf, b3 = cosf, b4 = sinf, b5 = sinf=cos u, b6 = cosf=cos u. In bond graph form, equation (13) is represented as a TF junction structure as shown in Figure 11.16,17,21,47,49,50 Rigid body dynamics in a multi-body system is generally modeled in the body-fixed reference frame aligned with the principal axes so that the Euler equations do not contain product of inertia terms. However, constraints between two or more rigid bodies with different body-fixed frames are implemented by transforming all motions to a common reference frame, that is, the inertial frame.47 The constraint forces and moments are then transformed back to the respective body-fixed frames",
            " In terms of scalar components taken along body-fixed Cartesian frame with unit vectors i\u0302, j\u0302, and k\u0302, the velocities turn out to be16,21,49 _xm, bi\u0302+ _ym, bj\u0302+ _zm, bk\u0302= _xbi\u0302+ _ybj\u0302+ _zbk\u0302 + vxbi\u0302+vybj\u0302+vzbk\u0302 3 xmi\u0302+ ymj\u0302+ zmk\u0302 \u00f027\u00de _xm, b = _xb + zmvyb ymvzb \u00f028\u00de _ym, b = _yb + xmvzb zmvxb \u00f029\u00de _zm, b = _zb + ymvxb xmvyb \u00f030\u00de Equations (28)\u2013(30) can be represented in BG form, in the support beam model as shown in Figure 15. Figure 15 shows one point on the rigid body. Four such points will be modeled in its full model. These are for the tail rotor and main rotor mount points, the pivot point and the counterweight support point. The self-weight of the beam needs to be modeled in the inertial frame Z-axis. For this purpose, a CTF from body-fixed frame to inertial frame through CTFBI (see Figure 12) is implemented in the bottom left part of Figure 16. The CTF requires Euler angles and hence, the EATF (see Figure 11) is also modeled in the top part of Figure 16. In Figure 17, four fixation points are modeled as indicated within the rectangular boxes. The x, y and z variables appearing in the various TF moduli indicate the positions from the CG of the TRMS beam and the subscript to those coordinates indicates the point, for example, p (pivot), m (main rotor and shield), t (tail motor and shield CG), and c (counterweight CG). In this particular application, the articulated joint between the beam and support column restricts the beam rotation to the inertial z-axis and body-fixed yaxis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002220_s11771-014-2071-8-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002220_s11771-014-2071-8-Figure1-1.png",
        "caption": "Fig. 1 Coordinate systems of an airship",
        "texts": [
            " Introduction Lighter-than-air vehicles have a wide range of applications, ranging from advertising, aerial photography and aerial inspection platforms, with a very important application area in environmental, biodiversity, and climatological research and monitoring. In recent years, the study for autonomous airships has become a hot research topic [1\u22122] because they exhibit a number of characteristics that outperform satellite, airplanes and helicopters in tasks that require autonomy, safety, low speed flight and low power to remain aloft (the coordinate system is depicted in Fig. 1). Among the research fields of airship, the attitude and position control is one of the most important component technologies. Many important results have been reported on airship control in the past years. For example, WANG and SHAN [3] studied the tracking control of airship based on the Lyapunov method. A backstepping controller was designed for path-tracking of an underactuated autonomous airship [4]. BEJI and ABICHOU [5] studied tracking control of a blimp by using a combined integrator backstepping approach and Lyapunov theory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001435_iecon.2013.6699639-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001435_iecon.2013.6699639-Figure11-1.png",
        "caption": "Fig. 11. For the case of two adjacent broken bars: a) Magnetic flux density; b) Current density distribution.",
        "texts": [
            " It can be observed from Fig. 10(a) that the radial forces are asymmetrical in this case, leading to a net radial force affecting the rotor shaft and, eventually, the bearings. Moreover, the tangential forces are no longer symmetric under the machine poles as shown in Fig. 10(b). The negative tangential forces cause a decrease in the average torque compared to the healthy case. Moreover, distributions for the magnetic flux density and the current density, for the case of two adjacent broken bars, are shown in Fig. 11. The tangential and radial forces are shown in Fig. 12. It can be observed from Fig. 12(a) that the asymmetry in the radial forces has increased in comparison to the case of one broken bar. Such increase in the asymmetry would negatively affect the bearings life time. Increased asymmetry of the tangential forces, as illustrated by Fig. 12(b), would result in increased torque and speed variations. The negative tangential forces cause a decrease in the average torque compared to the healthy case. It can be seen from the magnetic flux density distributions, shown in Fig. 7(a), Fig. 9(a) and Fig. 11(a) that the distribution in the faulty cases are asymmetric compared to the healthy case. This paper presents an efficient model to analyze broken bar fault conditions for induction motors using a FEM approach coupled with an ABC transient model. The pattern of the asymmetry in bar currents resulted from various broken bar faults can be deduced and analyzed. Effects on the motor torque and speed can be observed. Radial and tangential force asymmetries resulting from broken bars conditions can be assessed using the proposed approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002686_1.5062875-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002686_1.5062875-Figure2-1.png",
        "caption": "Figure 2: SLM test setup",
        "texts": [
            " In addition to the different pre-processing Selective Laser Micro Melting means having a laser spot and a part resolution < 100 \u00b5m and using powders < 50 \u00b5m. Due to the use of fine powders, the particles tend to agglomerate and special powder deposition systems are required. A special SL\u00b5M Machine system was developed to manufacture micro parts within the SLM process. The SLM test setup was built as a modular machine system having the capability to change optics, scanner, laser source and powder deposition mechanisms (see Figure 2). To enable SL\u00b5M the setup is equipped with a 50 W fiber laser and a 100 mm telecentric f-theta-lens reaching a laser spot diameter of 19.4 \u00b5m. The laser scanner is controlled by the commercial software tool SAM 3D (Scaps GmbH, Germany) which communicates with the machine control software developed with LabView (National Instruments, TX, USA) via TCP/IP. SAM 3D allows importing 3D models in the STL-format as well as slice data in CLI-format in ASCII or binary syntax. The CLI-data written in ASCII are of great importance for micro processing since every layer can be coded by the user himself using an easy to learn syntax"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002780_j.optlaseng.2020.106039-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002780_j.optlaseng.2020.106039-Figure8-1.png",
        "caption": "Fig. 8. (a) Temporal mode composed with standard PCA analysis. (b) Segmented defects where green denotes false defects and red positive defects (c) Temporal mode composed with weighted PCA analsyis (d) Segmented defects in red.",
        "texts": [
            "2. Bicycle frame Fig. 7 (a) shows an infrared image of a prototype bicycle part. At wo locations, the object is hit with an impact hammer (12 J). These l i r F p t a l m d 6 c m p f c t o a o r F \u201d q \u201d S ( w s ( D t C o w t a C P s A o ( R [ ocations are indicated in Fig. 7 (b). The CAD mapping result is visible n Fig. 7 (c) and the quality map in Fig. 7 (d). From the complete heat response, the standard principal thermogaphy analysis is performed. The third temporal mode is shown in ig. 8 (a). Fig. 8 (b) shows the segmented defects. Green areas are false ositives detected with this method. These areas are also areas where he quality Q is below 0.5. Note that an additional image-open filter (imge erosion followed by image dilation) is used to only highlight regions arger than 2 pixels. In Fig. 8 (c), the second temporal mode calculated with the weighted ethod is shown. Fig. 8 (d) shows the segmented defects in red. These efects correspond to the impacted locations. . Conclusion Information on the alignment of an infrared image with a CAD file an be used to predict measurement artefacts in the form of a quality ap. These predictions, together with a weighted version of the princi- al component thermography analysis, can be used to detect defects. In uture research, the calculated position of the camera can also be used to ope with the directional behaviour of the emissivity when calculating emperatures out of the measured data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.75-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.75-1.png",
        "caption": "Figure 4.75 Rotary motion; gearhead coupled",
        "texts": [
            "4 Results using Excel program in Table 4.3 for first motor selection Example in 4.9.1 1 2 BA N is the gear ratio of the gearhead. For most applications, this ratio is greater than 1 in order to create a condition in which the motor is operating at as high an efficiency as possible. For gearheads in which N is 10 or greater, the load inertia reflected back to the motor will often be relatively small, in which case a good initial approximation for the total inertia seen by the motor is JM + JGH. See Figure 4.75. System Data: JM = motor inertia (g cm s2) JL = load inertia (g cm s2) JGH = gearhead inertia (g cm s2) JS = shaft inertia (g cm s2) TL = load torque (g cm) TF = friction torque (g cm)(seen at gearhead input) System Design 235 Gearhead Efficiency (e): Spur or Bevel = 0.90; Worm = 0.65; Planetary = 0.95 Motion: Position: \u03b8M = N\u03b8L (rad) Velocity: \u03b8 \u2032 M = N\u03b8 \u2032 L (rad s \u22121) Acc/Dec: \u03b8 \u2032\u2032 M = N\u03b8 \u2032\u2032 L (rad s \u22122) At Motor: JT (total inertia) = JM + JL/N 2 + JGH (4.76) (theoretical, assuming the gearhead efficiency is 100%) In order to account for the actual gearhead efficiency, the torques are calculated as follows: TACC = (JM + JGH)\u03b8 \u2032\u2032 M + (JL/N 2e)\u03b8 \u2032\u2032 M \u00b1 TL/Ne + TF (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure1.18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure1.18-1.png",
        "caption": "FIGURE 1.18 A common conductor\u2013dielectric structure.",
        "texts": [
            " On the other hand, we must admit that these examples were not chosen only because they provide compact illustrative examples but because it is almost impossible to solve these equations formally for anything except very simple, highly symmetric, structures. Classical texts are replete with solution techniques and problems that have been solved.8 These are important in that they extend our knowledge of the mathematics and the physics of electrostatics. On the other hand, most situations that commonly arise in practice are either totally unsolvable by any of these techniques or would require months of analysis to produce a solution. As an example of this situation, consider Figure 1.18. We begin with a strip of metal on the top of a slab of dielectric that is fully metalized on its bottom face. For reasons that will be discussed in Chapter 2, we are interested in the capacitance between these two electrodes and the peak electric field (as a function of the applied voltage). This fairly straightforward situation is already a very difficult problem to attack analytically. Now let\u2019s add a narrow upper conductor strip branching off the original center conductor. What has happened to the capacitance and the peak electric field"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000716_ecce.2015.7309909-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000716_ecce.2015.7309909-Figure6-1.png",
        "caption": "Fig. 6: Cross section of eccentric internal rotor REL machine.",
        "texts": [
            " The rotor scalar magnetic potential is computed in [19]. It has a value only for the rotor islands as U (w) r = 2awRs \u2211 \u03c5 \u2212Ksn (np)2 cos(\u03bbnw) sin(np\u03b8b) (11) where aw and \u03bb\u03c5w are given by: aw = Rstb g\u0304wlb ( 1 + Rstb g\u0304wlb 2\u03b8b ) \u22121 (12) \u03bbnw = (2w \u2212 1)n\u03c0 2 + (n\u2212 1)p\u03b8m \u2212 \u03b1e i (13) and tb and lb are the thickness and length of the flux-barrier, respectively. The computation of the air gap flux density is carried out splitting the circumference of the rotor into 8 regions from \u03b8r = 0 to 2\u03c0 as shown in Fig. 6. Then, the average air gap length (10) is computed for each region. The air gap flux density Bg(\u03b8r) can be computed all around the rotor surface. Us(\u03b8r) and Ur(\u03b8r) vary with the rotor position so that the air gap flux density Bg(\u03b8r) can be computed for any rotor position. Details are in [19]. For both analytical models, the electromagnetic pressure on the rotor surface is given by: pm(\u03b8r) = B2 r (\u03b8r) 2\u03bc\u25e6 (14) The radial force on the rotor can be split into Fx and Fy in x-axis and y-axis directions. According to Fig. 3 and Fig. 6, the force along the stationary reference frame are: Fx = \u222b 2\u03c0 0 pm(\u03b8r)RsLstk cos(\u03b8r)d\u03b8r (15) and Fy = \u222b 2\u03c0 0 pm(\u03b8r)RsLstk sin(\u03b8r)d\u03b8r (16) In case of non-uniform displacement of rotor axis from the stator axis as shown in Fig. 1 (b) and Fig. 1 (c), the symmetric axis of the rotor is split into a finite number of slides ns. Then, the previous analytical models are applied considering each slide stack length equal to Lstk/ns. The total force in x axis and y axis directions are given by: Fx = ns\u2211 n=1 Fxn and Fy = ns\u2211 n=1 Fyn (17) where Fxn and Fxn are the force components acting on the n\u2212 th slide"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002239_1350650115592919-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002239_1350650115592919-Figure1-1.png",
        "caption": "Figure 1. Investigated tilting-pad bearing and pad geometry.",
        "texts": [
            "comDownloaded from not have the technical capabilities by the means of test rigs to reliably find new effects. In this paper, an advanced, axially concave pad profile is proposed which is capable of compensating for the malicious thermomechanical effects of large turbine tilting-pad bearings. The investigations were carried out on a five-tiltingpad bearing for power generation applications on the test rig for large turbine bearings at the RuhrUniversity Bochum. A schematic illustration of the bearing and pad geometry is given in Figure 1. The bearing is mounted in load between pivot position. Pads and bearing shell both feature a radius in axial direction which leads to a point contact pivot. Due to Hertzian stress, the pivot becomes elliptical. This setup provides a relatively high capability of tilting in axial direction, and thus prevents larger influences of misalignment between bearing and shaft. Due to the resulting small elliptical area of contact between pad and liner, this pivot is very flexible. The backs of the pads are directly in contact with the liner in the middle plane of the bearing, as the pads are fixed by holding pins on the leading edge and spray bars on the trailing edge"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003214_jmems.2020.3005090-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003214_jmems.2020.3005090-Figure1-1.png",
        "caption": "Fig. 1. Fabrication process and released devices: (a) MLS Fabrication Process. Manual Steps in orange. (b) Image of bronze switches on build platform after printing. (c) Batch fabricated devices in frames after EDM wire release.",
        "texts": [
            " Device microfabrication was performed using micro laser sintering, a precision form of direct metal laser sintering (DMLS), where components are additively manufactured by sintering, or melting, a metal powder in a layer-by-layer fashion using a laser. This technology utilizes smaller diameter metal powders <10\u03bcm and higher-precision optics to enable finer component features [23]. Herein, a DMP50GP MLS 3D printer (3D Microprint GmbH) using 5\u03bcm thick layers, and a laser spot size of 30\u03bcm at 40W CW power in an argon environment were employed. Bronze (90% Cu - 10% Sn) powder (D90 6\u03bcm) was procured from Dongguan Hyper Powder Technology Co. Ltd. Fabrication proceeds as follows and is detailed in [12] and is illustrated in Figure 1(a): First, after device design in computer aided drafting (CAD) software, the 3D model is processed into a laser manufacturing plan using commercial slicing software (AutoFab, Renishaw plc.). The bronze powder is filtered using a series of vibratory mesh sieves (63\u03bcm, 36\u03bcm, 25\u03bcm) in an argon environment to remove any large contaminants. Following this, the metal powder and manufacturing plan are loaded into the 3D printer. A 57mm diameter stainless steel platen is loaded into the system to serve as a build platform, with the components fusing onto the plate during manufacturing. Device fabrication occurs once the system is activated, and then for each device layer, 5\u03bcm of powder is applied on top of the build platform via a doctor blade. Next, the shape of the device cross-section is melted via laser on the powder, forming solid metal. After patterning, the build platform lowers by 5\u03bcm to receive another layer of unmelted powder, and then process repeats until the entire component is fabricated as in Figure 1(b). The fabrication process is described graphically using Figure 2 and is performed such that the top plane face is layered first and the preceding bottom layers are added successively. This will aid creating electrodes Authorized licensed use limited to: Carleton University. Downloaded on July 25,2020 at 21:55:21 UTC from IEEE Xplore. Restrictions apply. with different heights as per the process design and prevent overhangs. After fabrication, the build plate is removed from the 3D printer and is sonicated for 30 minutes to remove any latent metal powder. Wire cut EDM, using a 400\u03bcm diameter wire is then used to detach the printed components from the build platform. The components are then rinsed in isopropyl alcohol and dried before further processing. Figure 1(c) shows the released devices. Scanning electron microscopy was used to characterize the as-printed components, Figure 3. The first step towards the transfer of the MEMS component on a substrate is cleaning. A soda lime glass slide substrate was used for the transfer. Before the transfer, the glass surface was cleaned using lab-grade cleaning paper. Figure 4 shows the CAD design of the MEMS switch to be transferred. The green surfaces of the device are the anchor points. Figure 4e and 4f shows the side and cross-sectional view of the device with green anchor points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003329_aim43001.2020.9158972-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003329_aim43001.2020.9158972-Figure8-1.png",
        "caption": "Figure 8. Definition of rotation",
        "texts": [
            " 7, the shape of the graph of velocity change is maintained even if slip occurs. In addition, the component of acceleration due to gravity is added to the acceleration measured by the sensor when the moving distance is calculated in three-dimensional space. By subtracting the acceleration vector of the component force of gravity (ax, ay, az) from each acceleration component (ASensor x, ASensor y, ASensor z) obtained from the sensor, the acceleration vector of the component force of gravity can be defined (Ax, Ay, Az) [(1)]. As shown in Fig. 8, (ax, ay, az) can be calculated from (2) when the angle from the initial position of the observed sensor is represented by and along the x-, y-, and z-axes. Where C cos and S sin The bending pipe is identified from the change in angular velocity and attitude angle as measured by the IMU. Fig. 9 shows a schematic diagram of the pipeline. As shown in Fig. 9, straight pipes are connected before and after the bending pipe in the actual pipeline. Fig. 10 shows the measured values of the angular velocity when the sensor sequentially passes through a bending pipe sandwiched between straight pipes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003628_mra.2020.3024395-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003628_mra.2020.3024395-Figure3-1.png",
        "caption": "Figure 3. (a) The Hapkit components, including laser-cut acrylic pieces, 3D-printed parts, custom PCB, motor, various nuts and screws, power supply, micro-USB cable, and tools needed for assembly. (b) The simple assembly process features laser-etched numbers on each acrylic piece along with a series of tabs and slots to enable easy connection of corresponding pieces. (c) The handle stop is designed to prevent the sector pulley from continuous rotation that would otherwise cause the handle to collide with the spinning motor and break.",
        "texts": [
            " A fully assembled, 1-DoF, low-cost, open source haptic device, Hapkit, used in hands-on laboratories in an online course. Design features that ensure students could assemble, program, and interact with this device in an online learning environment are highlighted. MR: magnetoresistive; SD: Secure Digital. Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 14:34:18 UTC from IEEE Xplore. Restrictions apply. Unlike previous Haptic Paddles, Hapkit was intended for students to build completely on their own using a kit, shown in Figure 3(a), which includes all of the components and tools necessary for assembly. The main body of Hapkit is made of laser-cut acrylic pieces. To ease assembly, a number is etched onto each of these pieces [Figure 3(b)] as well as a series of tabs and corresponding holes on parts meant to connect. Glue, three hex keys, and a small flathead screwdriver are the only tools required. A list of parts, tools, and cost are provided in the Hapkit 1.0 section of the project website, http://hapkit.stanford.edu. The original Haptic Paddle, along with several successors at various universities, used a capstan drive mechanism [11], [15]. This design requires winding a thin cable around a capstan (which is attached to the motor) several times and then properly tensioning the cable enough to transmit forces from the motor to the sector pulley",
            " Alternatively, a friction drive, while adding a considerable amount of inherent damping to the system, is far less complicated to assemble. After initially attaching a piece of neoprene rubber to the bottom edge of the sector pulley, the amount of friction can be easily adjusted by raising or lowering the bar shown in Figure 1. If the device becomes unstable, students simply rotate the sector pulley until it is back in contact with the drive wheel. An additional design feature that helps students easily reset the device if it goes unstable is the handle stop, shown in Figure 3(c). If the motor forces the sector pulley hard enough in one direction so that it loses contact with the drive wheel, the handle stop hits the top of Hapkit, preventing the sector pulley from continuously spinning. This feature prevents Hapkit from self-damage of the handle, since it cannot collide with the spinning drive wheel, as well as from wearing of the neoprene, which does not contact the spinning drive wheel. Another important design choice was the size of the drive wheel. Because the radius of the drive wheel, radius of the sector pulley, and length of the handle determine the force felt by the user, as shown in (2), it was important to have a smaller drive wheel to increase the force felt"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003147_j.engstruct.2020.110892-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003147_j.engstruct.2020.110892-Figure3-1.png",
        "caption": "Fig. 3. Loading Configuration of the IMT.",
        "texts": [
            " The interaction can then be quantified by comparing results from both cases. The IMT in the experiment has circular cross-section with the external radius of 50 mm. Its specimens are of different lengths to study the influence of slenderness ratio. They are fabricated by the authors with Kapton\u00ae membrane with the thickness of 0.075 mm. Such an IMT is popularly used as basic members of a solar sail in aerospace engineering as stated in Section 1. The IMT was subjected to both bending and axial compression as shown in Fig. 3 (where \u03bb, L and d respectively denote the slenderness ratio, length and diameter of the IMT with L = \u03bbd). Three kinds of constraints, i.e. fixed support1, simple support and the combination of both, were applied separately to the ends of the tube. Two transverse concentrated forces FH were applied at one-third lengths of the tube and the axial compressive force FV was applied at the top. The experiment was performed in the student dormitory of Beijing Jiaotong University. The axial force FV is applied first and fixed at the specified value in the whole loading period"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000760_icacci.2014.6968225-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000760_icacci.2014.6968225-Figure2-1.png",
        "caption": "Fig. 2: Angular misalignment between two shafts.",
        "texts": [
            " The 2X part is commonly larger than the IX, however its size relative to the IX is commonly determined by the sort and construction of the coupling [13] (see Figure 3 below). Once the angular or parallel misalignment is important, it will generate either high amplitude peaks at many harmonics (4X to 8X) or maybe lots of harmonics to a high frequency that is comparable to the mechanical looseness. 2) Angular misalignment: In this case the misalignment shafts meet at one purpose however they are not in parallel (Figure 2). 20 i4 International Conference on Advances in Computing, Communications and informatics (ICACCI) 2477 Angular misalignment is characterized by massive axial vibration that is out-of-phase, i.e. with the distinction 1800 over the coupling. In a typical case, large axial vibration is on each IX and 2X elements [13] (see Figure 3 below). However it's quite common that any of the elements IX, 2X or 3X dominates. These symptoms can also indicate the existence of the issues with the coupling. The wide angular misalignment could excite several harmonics of rotational frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003557_s40430-020-02663-1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003557_s40430-020-02663-1-Figure2-1.png",
        "caption": "Fig. 2 Equivalent process for hobbing helical non-circular gears",
        "texts": [
            " However, this formula is not suitable for calculating the additional rotation of the noncircular gear blank because the non-circular characteristic of the pitch curve cannot ensure that angle \u03bc will become equal to 90\u00b0 at each position, as is the case for a cylindrical gear. (2) When considering the influence of vz on \u03c9c, the latter\u2019s influence on vx is ignored. As shown by Eq.\u00a0(2), vx is mathematically related to vy; in case of a helical gear, vz will affect vy, affecting vx. Based on the existing research problems, an improvement of the hobbing mathematical model in case of helical noncircular gears is presented below. First, the influence of vz on vy must be clarified. The hobbing process is equivalent to the meshing process of the helical gear and rack (Fig.\u00a02). As shown in Fig.\u00a02, the hobbing process of the helical non-circular gears is equivalent to the meshing transmission process of the helical rack and gear. Each moment of the hobbing process is equivalent to the momentary meshing state of the infinite thin rack and the non-circular spur (5)\u0394 c = vz tan r gear. Point Q1 is located on the centerline of the equivalent rack in the M\u2013M section, line segment P1P2 is located on the pitch plane of the equivalent helical rack, and the angle between P1P2 and the gear rotation axis is the helical angle \u03b2. When the meshing state of the hobbing process moves from the M\u2013M section to \u2206z and subsequently to the N\u2013N section, the hob is assumed to be exactly rotated by an integer number of circles. Then, at this moment, point Q1 will correspond to point Q2 located directly below instead of point Q3 on line segment P1P2. In case of the up cutting process of the right-handed non-circular gear shown in Fig.\u00a02, the movement speed of the rack of the equivalent helical rack in the N\u2013N section will be accelerated because of the movement of the hob in the vertical direction. During the helical non-circular gear hobbing process, the moving speed of the projection rack in the y-axis direction is vy*. In this case, where mn is the normal modulus and Kz is the sign coefficient. Kz is considered to be 1 when the non-circular gear is left-handed and processed by down-milling or right-handed and processed by up-milling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure14.3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure14.3-1.png",
        "caption": "Fig. 14.3 Two body abrasive wear testing machine",
        "texts": [
            " Working pressure for all the experiments is kept at 6 kg/cm2 (0.589 MPa). Anyonewhile investigating friction andwear ofmetals is facedwith a large number of variables to control. One of the first choices tomake is the selection of a suitable wear testing machine. It has become axiomatic that the best wear test is one that closely approximates the actual conditions encountered in service. However, here high stress abrasive wear tests are done using SUGA (NUSI, JAPAN) abrasion testing machine. Schematic view of the machine is shown in Fig. 14.3. Two body abrasive wear tests were conducted on 45 \u00d7 45 \u00d7 4.5 mm3 size specimens using Suga Abrasive Tester (Model: NUSI, Japan). A schematic diagram of the test apparatus is shown in Fig. 14.3. The equipment consisted of a stage with a locking arrangement for holding the specimens in position against a 50mmdiameter, 12 mm thick metallic disc. Emery paper embedded with silicon carbide particle was cut into sizes and fixed on a wheel with the help of double sided tape to serve as the abrasive medium. The locking arrangement was used to fix the specimen with the abrasive medium. The stage was attached with a motor; the latter imparted to and fro motion to the specimens along with facilities to display the stage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002935_b978-0-323-66164-5.00006-4-Figure6.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002935_b978-0-323-66164-5.00006-4-Figure6.1-1.png",
        "caption": "FIGURE 6.1",
        "texts": [
            " It is defined as the process of joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies. CHAPTER 853D Printing: Applications in Medicine and Surgery. https://doi.org/10.1016/B978-0-323-66164-5.00006-4 Copyright \u00a9 2020 Elsevier Inc. All rights reserved. The start of process is a digital file of the item that can be created using a CAD tool, or digitized if already existing (by a scanner or tomography). Having the design of a product is the first step for printing it (making it additively), that can be made anywhere in the world (providing a suitable machine and the raw material). Fig. 6.1 shows the process of building the part from the 3D digital model. Under the umbrella of AM there are many processes. ASTM groups them in seven types (Fig. 6.2): 1) Binder jettingdAM process where a liquid bonding agent is deposited to join powdered materials together. 2) Direct energy deposition (direct manufacturing)dAM process where thermal energy fuses or melts materials together as they are added. 3) Material extrusion (fused deposition modeling)dAM process that allows for depositing material via a nozzle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000869_20140824-6-za-1003.01890-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000869_20140824-6-za-1003.01890-Figure7-1.png",
        "caption": "Fig. 7. Drad and Lift Forces of Rotated Flying Ball.",
        "texts": [
            " Then, the aerodynamics model of the ball is given by (Nonomura et al. (2010)) mp\u0308 = \u2212mg \u2212 1 2 \u03c1SbCD(p\u0307,\u03c9)||p\u0307\u2225p\u0307+ \u03c1VbCM (p\u0307,\u03c9)\u03c9 \u00d7 p\u0307, (1) where Sb := \u03c0r2 is the ball cross section area, Vb := 4 3\u03c0r 3 is the ball volume, r = 0.02 [m] is the ball radius,m = 0.0027 [kg] is the ball mass, \u03c1 = 1.184 [kg/m3] is the air density and and g := [0 0 g]T with g = 9.8 [m/s2] is the acceleration of gravity. In the right hand of (1), the second and third terms represent the drag and lift effects with their coefficients CD and CM as shown in Fig. 7. The coefficients depend on the directions of the rotational velocity \u03c9, i.e. the top, back and side spins, which are described in Fig. 8. In the left figure, p\u0307xy := [p\u0307x p\u0307y 0] \u2208 R3 represents the x and y components of the ball velocity in the reference frame \u03a3B. The frame \u03a3Ball is defined such that the x\u2032- and z\u2032-axes are along the direction of p\u0307xy and the z-axis of \u03a3B . The orientation of \u03a3Ball relative to \u03a3B is represented by the rotational matrix RBall(\u03c8) := [ cos\u03c8 \u2212 sin\u03c8 0 sin\u03c8 cos\u03c8 0 0 0 1 ] , (2) where \u03c8 is the angle between the x- and x\u2032- axes of the frames defined by cos\u03c8 := p\u0307x\u221a p\u03072x + p\u03072y , cos\u03c8 := p\u0307y\u221a p\u03072x + p\u03072y "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002861_tasc.2020.2977592-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002861_tasc.2020.2977592-Figure3-1.png",
        "caption": "Fig. 3. Distribution of horizontal illumination at the CF, Minimum 336 lx and Maximum 894 lx.",
        "texts": [
            ", 4, 6 and 8 LD units) were tested to find the optimum number of units that ensure that the ISO and EU standards illumination requirements are satisfied in the room. We found that eight units were the optimum for illumination, and we used this in our study; four and six light units in the room did not achieve the minimum illumination requirement (i.e., 300 lx [25]). The height of the work desks where the transmitters and receivers associated with the user equipment are placed was 1m. This horizontal plane was referred to as the \u201ccommunication floor\u201d, (CF), in Fig. 1(a) and (b). Fig. 3 shows the horizontal illumination distributions from the eight RGB-LD light units at the CF level. It is clear from this figure that there is sufficient illumination according to EU and ISO standards [25]. In optical wireless (OW) links including VLC system, intensity modulation with direct detection (IM/DD) is the preferred choice [28], [36]. IM can be simply defined as the instantaneous power of the optical carrier modulated by the signal. At the receiver side, DD is used to generate electrical current proportional to the instantaneous received optical power"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002971_j.matpr.2020.03.552-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002971_j.matpr.2020.03.552-Figure1-1.png",
        "caption": "Fig. 1. Plate g",
        "texts": [
            " Moreover, the computational cost can be further improved by utilizing the commercial tool (ANSYS and ABAQUS) which are well accepted in the industries. Now, the present research aims to investigate the modal frequencies of Luffa cylindrica sponge fibre-reinforced polymer composite via a simulation model derived in ANSYS environment using the batch technique. The model validity has been checked with the experimental and available published numerical frequency data including the experimental properties. The geometry of the laminated composite plate as shown in Fig. 1 considered for the simulation study. The dimensions of any plate model developed in ANSYS environment using the corresponding length \u2018a\u2019, width \u2018b\u2019 and thickness \u2018h\u2019. The model input data has been introduced via ANSYS parametric design language code (APDL) for the batch case. The structural deformation model generally utilized the in-built first-order shear deformation theory (FSDT) in ANSYS [11] platform and conceded as: U\u0302\u00f0X1;X2;X3; t\u00de \u00bc U\u03020\u00f0X1;X2; t\u00de \u00fe X3/X1 \u00f0X1;X2; t\u00de V\u0302\u00f0X1;X2;X3; t\u00de \u00bc V\u03020\u00f0X1;X2; t\u00de \u00fe X3/X2 \u00f0X1;X2; t\u00de W\u0302\u00f0X1;X2;X3; t\u00de \u00bc W\u03020\u00f0X1;X2; t\u00de \u00fe X3/X3 \u00f0X1;X2; t\u00de \u00f01\u00de where,U\u0302, V\u0302 and W\u0302are the displacement of any point on the kth layer of the plate at time \u2018t\u2019 along theX1, X2 andX3 material coordinate axes whereas U\u03020, V\u03020 and W\u03020are the corresponding displacements of point on the mid-plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure5.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure5.1-1.png",
        "caption": "Fig. 5.1 Schematic representations of the abrasion wear",
        "texts": [
            " As per Gahr, wear is defined as the \u201cprogressive loss of substance from the operating surface of a body occurring as a result of relative motion at the surface\u201d (Zum Gahr 1987). Wear is classified into six types. These are mainly (Burwell 1957; Sharma et al. 2011): (1) Abrasive: Abrasive wear can be defined as a cutting or loss of mass when a hard surface slides or roll under pressure on another surface. In abrasionwear test, the sample can be slide against a rough counter face or abrasion by hard particles Abrasion may be two bodies or three bodies depending upon the number of the interacting surfaces. Figure 5.1 shows the schematic representations of the abrasion wear. (2) Adhesive: Adhesive wear can be defined as mutual affinity between the materials which causes material exchange between the two mating surfaces or the loss from either surface. Figure 5.2 shows the schematic representations of the adhesive wear. (3) Erosive: Erosivewear is amaterial removal processwhen a solid particles impact on a surface. This material removal process may be occurred by gas or liquid which flow with and without carrying solid particles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002965_02640414.2020.1754717-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002965_02640414.2020.1754717-Figure9-1.png",
        "caption": "Figure 9. A superior plane view of a seam azimuth angle of 30\u00b0, a measure of the angle created by the seam plane (dotted line) and the ball\u2019s direction of travel.",
        "texts": [
            " Ball velocity and angular velocity were computed using the magnitude of the respective vectors from data collected 30 frames post-BR (Spratford et al., 2017). Seam azimuth angle was determined using the mean of three frames post-BR due to potential variability in seam position (Spratford et al., 2017), and therefore, was treated as a discrete variable and not filtered (Linthorne & Patel, 2011). Seam azimuth angle increased positively in an anticlockwise direction and negatively in a clockwise direction (Figure 9). Seam stability was calculated and expressed as a percentage (Spratford et al., 2017). A value of 100% represented perfect stability, where the ball\u2019s seam and axis of rotation were perpendicular. A value of 0% represented perfect instability, where the ball\u2019s seam and axis of rotation were coincident, commonly referred to as a crossseam delivery. The swing of each delivery was determined using ball pitch location from camera images and initial flight trajectory captured by the Vicon MX system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003021_b978-0-12-418691-0.00003-4-Figure3.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003021_b978-0-12-418691-0.00003-4-Figure3.2-1.png",
        "caption": "FIGURE 3.2 Schematic representation of an oscillatory motion of amplitude C, angular frequency \u03c9, and initial phase \u03c8.",
        "texts": [
            " Depending on the presence or absence of energy dissipation mechanisms, we may have damped forced vibration or undamped forced vibration. Again, the undamped forced vibration is easier to analyze, but the damped forced vibration is more representative of actual phenomena. An oscillatory motion can be defined by the formula u\u00f0t\u00de5C cos\u00f0\u03c9t1\u03c8\u00de (2) STRUCTURAL HEALTH MONITORING WITH PIEZOELECTRIC WAFER ACTIVE SENSORS where C is the amplitude measured in length units, \u03c9 is the angular frequency measured in radians per second (rad/s), and \u03c8 is the initial phase, measured in radians (Figure 3.2). The frequency, f, which is measured in cycles per second (c/s) or Hz, is related to the angular frequency by the formula f 5 1 2\u03c0 \u03c9 (3) The period, \u03c4, measured in seconds, is related to the frequency, f, by the formula \u03c45 1 f (4) In phasor notation, the vibrational motion is represented by the phasor C+\u03c8, where C is the magnitude and \u03c8 is the phase angle. Recall Euler identity (ref. [1], p. 24) ei\u03b1 5 cos\u03b11 isin\u03b1, \u03b1A\u211d (5) Using Euler identity given by Eq. (5), we can view the cosine function as the real part of the complex exponential function, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure2-1.png",
        "caption": "Fig. 2. Six-axis Cartesian-type CNC hypoid gear generator.",
        "texts": [
            "012 where Pleas (201 a12 \u00bc \u2212 cos j\u2212q\u00f0 \u00de cos\u03b3m cos\u03bcg \u00fe cosi cos\u03b3m sin j\u2212q\u00f0 \u00de \u00fe sini sin\u03b3m sin\u03bcg a13 \u00bc \u2212 cos\u03b3m sini sin j\u2212q\u00f0 \u00de \u00fe cosi sin\u03b3m a14 \u00bc \u2212\u0394A\u00fe SR cosq cos\u03b3m\u2212\u0394B sin\u03b3m a21 \u00bc cos\u03bcg\u00f0 cosi cos j\u2212q\u00f0 \u00de cos\u03d51 \u00fe \u00f0 cos\u03b3m sini\u2010 cosi sin j\u2212q\u00f0 \u00de sin\u03b3m\u00de sin\u03d51\u00de\u2212 cos\u03d51 sin j\u2212q\u00f0 \u00de \u00fe cos j\u2212q\u00f0 \u00de sin\u03b3m sin\u03d51 sin\u03bcg a22 \u00bc \u2010 cos\u03bcg cos\u03d51 sin j\u2212q\u00f0 \u00de \u00fe cos j\u2212q\u00f0 \u00de sin\u03b3m sin\u03d51\u00f0 \u00de\u2212 cos\u03b3m sini\u2010 cosi sin j\u2212q\u00f0 \u00de sin\u03b3m\u00f0 \u00de sin\u03d51 \u2212 cos\u03d51 sin j\u2212q\u00f0 \u00de \u00fe cos j\u2212q\u00f0 \u00de sin\u03b3m sin\u03d51\u00f0 \u00de sin\u03bcg a23 \u00bc cos j\u2212q\u00f0 \u00de cos\u03d51 sini\u2212\u00f0 cosi cos\u03b3m \u00fe sini sin j\u2212q\u00f0 \u00de sin\u03b3m\u00de sin\u03d51 a24 \u00bc SR cosq sin\u03b3m sin\u03d51 \u00fe cos\u03d51 Em\u2212SR sinq\u00f0 \u00de \u00fe \u0394B cos\u03b3m sin\u03d51 a31 \u00bc cos\u03bcg\u00f0\u2212 cos\u03b3m cos\u03d51 sini\u00fe cosi\u00f0 cos\u03d51 sin j\u2212q\u00f0 \u00de sin\u03b3m \u00fe cos j\u2212q\u00f0 \u00de sin\u03d51\u00de\u00de \u00fe cos j\u2212q\u00f0 \u00de cos\u03d51 sin\u03b3m\u2212 sin j\u2212q\u00f0 \u00de sin\u03d51\u00de sin\u03bcg a32 \u00bc cos\u03b3m cos\u03d51 sini sin\u03bcg\u2212 sin j\u2212q\u00f0 \u00de\u00f0 cos\u03bcg sin\u03d51 \u00fe cosi cos\u03d51 sin\u03b3m sin\u03bcg\u00de \u00fe cos j\u2212q\u00f0 \u00de\u00f0 cos\u03bcg cos\u03d51\u2212 cosi sin\u03d51 sin\u03bcg a33 \u00bc cos\u03d51 cosi cos\u03b3m \u00fe sini sin j\u2212q\u00f0 \u00de sin\u03b3m\u00f0 \u00de \u00fe cos j\u2212q\u00f0 \u00de sini sin\u03d51 a34 \u00bc sin\u03d51 Em\u2212SR sinq\u00f0 \u00de\u2212 cos\u03d51 SR cosq sin\u03b3m \u00fe \u0394B cos\u03b3m\u00f0 \u00de A mathematical model of a six-axis Cartesian-type CNC hypoid gear generator is presented in [2]. Fig. 2 shows the coordinate systems for the six-axis Cartesian-type CNC machine. The coordinate systems Sts and S1 s are rigidly connected to the cutter and the workpiece, respectively, and the transformation matrices from St s to S1 s define the spatial relationship between the cutter and the workpiece in the coordinate system S1 s , as shown in Eq. (2). Ms pt \u00bc Ms pdM s deM s emM s mhM s ht \u00bc S11 S12 S13 S14 S21 S22 S23 S24 S31 S32 S33 S34 0 0 0 1 2 664 3 775 \u00f02\u00de S11 \u00bc cos\u03d5S B cos\u03d5 S C S12 \u00bc cos\u03d5S B cos\u03d5 S C S13 \u00bc sin\u03d5S B S14 \u00bc Mt \u00fe Ct \u00fe Tt \u00fe cos\u03d5S B\u0394 S X\u2212 sin\u03d5S C\u0394 S Z e cite this article as: S",
            " In the experiments, 40% of machining time was reduced by the proposed feed rates under the same cutting torque; the machining time is also helpful to protect the tools and machine from excessive cutting torque. The gear can be manufactured with higher efficiency on the six-axis CNC hypoid-gear generator with the proposed CMRR method. It is also applied in the grinding process. Nomenclature \u03d5c cradle angle \u03d51 workpiece rotation angle \u03b3m machine root angle \u03bcg cutter rotation angle i tilt angle j swivel angle q cradle rotation angle \u0394A increment of machine center to back \u0394B sliding base feed setting SR radial setting Em vertical offset \u03d5A,B,C = rotational axes of the six-axis machine, as shown in Fig. 2 Mt = workpiece mounting distance, as shown in Fig. 2 Ct = arbor distance, as shown in Fig. 2 Tt = distance between rotational center of machine root angle and workpiece mounting plane, as shown in Fig. 2 Cd = distance between the pitch apex of work gear and the origin Oc, as shown in Fig. 2 \u0394A,B,C = axis displacement of the six-axis machine, as shown in Fig. 2 Et = offset between rotational axis of machine root angle and workpiece spindle rm mean radius \u03b3 pitch cone angle \u03b3o face cone angle \u03b3r root cone angle Lo face apex beyond crossing point Lr root apex beyond crossing point \u03b1co outside blade angle \u03b1ci inside blade angle rcm cutter radius wc cutter point width hco cutter tooth height \u0394\u03d5 interval angle s 6-axis CNC machine subscript a workpiece rotation axes b workpiece spindle pivot axes c cutter rotation axes t cutter head coordinate system 1 workpiece coordinate system [1] Z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure7-1.png",
        "caption": "Fig. 7 Schematic Diagram of the Mechanism Kinematics",
        "texts": [
            " Therefore, the flexible cushioning element is designed and installed between the ropes and ball-screw pairs to tolerate the inconsistent rope lengths, as shown in Fig. 6. To meet both the requirements of tolerance and accuracy, the stiffness of the flexible cushioning element has been carefully designed and verified. Due to numerous motion joints, To describe and analyze the posture of the snake manipulator prototype, it is necessary to establish an improved kinematic model. The prototype is described using the D-H coordinate system, as shown in Fig. 7. In the figure, Cn is the center of the universal joint of the segment n. Pn(n = 1, 2, \u00b7 \u00b7 \u00b7 5) is the position vector of Cn. According to the mechanical configuration, the D-H parameters can be determined, as shown in Table 1. In Fig. 7, the transformation matrix between adjacent D-H coordinates is: i\u22121Ti = \u23a1 \u23a2\u23a2\u23a3 ci \u2212c\u03b1isi s\u03b1isi aci si c\u03b1ici \u2212s\u03b1ici aisi 0 s\u03b1i c\u03b1i di 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6 ci = cosqi, si = sinqi, c\u03b1 = cos\u03b1i, s\u03b1 = sin\u03b1i (1) The end posture can be obtained from (1) is: TI = T0 \u00b7 0T1 \u00b7 1T2 \u00b7 \u00b7 \u00b7 I\u22121TI = [ RI pI 01\u00d73 1 ] (2) Where, I = 10 is the quantity of the D-H coordinates; pI =[ x y z ]T represents the end position vector. RI represents Table 1 D-H parameters Linki 1 2 3 4 5 6 7 8 9 10 qi \u03b81 \u03b82 \u03b83 \u03b84 \u03b85 \u03b86 \u03b87 \u03b88 \u03b89 \u03b810 di(mm) 0 0 0 0 0 0 0 0 0 0 a(mm) 0 110 0 110 0 110 0 110 0 50 \u03b1(\u25e6) -90 90 -90 90 -90 90 -90 90 -90 90 the attitude rotation matrix of the end coordinate system, and can be expressed by the the Euler angles \u03c6, \u03b8, \u03c8 : RI = \u23a1 \u23a3 cos \u03b8 cos\u03c6 sin\u03c8sin\u03b8 cos\u03c6 \u2212 cos\u03c8 sin\u03c6 cos \u03b8 sin\u03c6 sin\u03c8sin\u03b8 sin\u03c6 + cos\u03c8 cos\u03c6 \u2212 sin \u03b8 sin\u03c8 cos \u03b8 cos\u03c8 sin \u03b8 cos\u03c6 + sin\u03c8 sin\u03c6 cos\u03c8 sin \u03b8 sin\u03c6 \u2212 sin\u03c8 cos\u03c6 cos\u03c8 cos \u03b8 \u23a4 \u23a6 (3) The number of holes is the same as ropes, which is j = 1, 2, 3 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000351_s1446181118000019-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000351_s1446181118000019-Figure6-1.png",
        "caption": "Figure 6. Schematic diagram showing the three frames involved in the problem of a buckling ring. (a) The ungrown (attachment) radius, (b) the unstressed grown radius, (c) the current, buckled state.",
        "texts": [
            " Geometry Following the general notion of the decomposition of the deformation gradient in morphoelasticity [15] and applied to elastic rods as in the framework of morphorods [11], there are three frames we must consider for this problem: an initial, stress-free circular frame of radius A and parameterized by arc length, S0; a hypothetical grown but stress-free frame which remains circular but with radius \u03b3A and parameterized by arc length S ; and the current, stressed state, parameterized by its arc length, s, as indicated in Figure 6. The objective is to determine the smallest value of \u03b3 > 1 for which the current state admits a noncircular, buckled solution, and to determine the corresponding mode number for this buckled state. As we are working in an initially circular geometry, a polar coordinate system, {er, e\u03b8, ez} is an obvious choice, although note that the majority of the analysis requires only planar considerations. We also make use of a local, right-handed frame defined by the current (buckled) configuration, {d1, d2, d3}, where d1 is associated with the inward pointing normal, d2 with ez and d3 with the tangent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure15-1.png",
        "caption": "Fig. 15. The Singular configuration II.",
        "texts": [
            " Starting from this Singular configuration I ( Fig. 13 ), when the joint axes of R joints J R2 , J R4 , J R6 , and J R8 are always parallel and in the same directions during the movement, the mechanism makes a planar motion, which is called plane-motion Mode III. Its typical configurations are shown in Fig. 14 . When the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R1 and J R5 are parallel, and the joint axes of the R joints J R3 , J R7 are also parallel, as shown in Fig. 15 . This configuration is a special configuration of Mode I. Since the axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration II ( Fig. 15 ), when the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R2 and J R6 are in the same direction, the joint axes of R joints J R4 and J R8 are in the same direction, but the two groups directions of joint axes are opposite during the movement. Then the mechanism makes a planar motion, which is called plane-motion Mode IV. Its typical configurations are shown in Fig. 16 . When the axes of R joints J R2 and J R8 coincide, the axes of R joints J R4 and J R6 coincide, and the axes of R joints J R3 and J R7 coincide, the configuration is a special configuration of Mode II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001756_978-3-319-24502-7_10-Figure10.6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001756_978-3-319-24502-7_10-Figure10.6-1.png",
        "caption": "Fig. 10.6 A non-Brownian sphere settling with translational velocity V under gravity of field strength g near a vertical hard wall. The principle of kinematic reversibility in conjunction with the flow geometry leads to the conclusion that the sphere maintains a fixed distance from the wall while settling and rotating clockwise. See the text for details",
        "texts": [
            " When the rotation is reversed subsequently by the same number of turns, the original droplet is reconstituted up to a small amount of blurring originating from the irreversible residual Brownian motion of the dye particles. The length of the filament depends on the number of turns only, independently of the rate at which the inner cylinder is rotated. This nicely illustrates the earlier discussed instantaneity of Stokes flows. The kinematic reversibility in combination with specific symmetries puts general constraints on the motion of a microparticle in a viscous fluid. A classical example is a spherical rigid microparticle settling under gravity near a stationary vertical hard wall (see Fig. 10.6). While the particle is rotating clockwise during settling, owing to the larger wall-induced hydrodynamic friction on its semi-hemisphere facing the wall (see Sect. 10.5.2 for details), a question arises whether it will approach the wall or recede from it. Given that gravity acts vertically downwards parallel to the wall, assume for the time being that the sphere approaches the wall while settling (see Fig. 10.6a). Kinematic reversibility requires that once the direction of the motion-driving gravitational force is reversed, the Stokes flow pattern remains unchanged except for the directional reversal of the fluid elements motion, provided the translational and angular particle velocities are likewise reversed. According to Fig. 10.6b, this implies that the sphere sediments upwards while receding from the wall. On rotating Fig. 10.6b by 180\u00b0 around the horizontal symmetry axis line going through the sphere centre, Fig. 10.6c is obtained in conflict with Fig. 10.6a wherein the sphere had been assumed to approach the wall. A contradiction is avoided only if the sphere remains at a constant distance from the wall while settling, as in Fig. 10.6d. An analogous reasoning can be employed to show that in a Poiseuille channel flow, a non-Brownian microsphere translates along the flow streamline, without any cross-flow velocity component. As discussed in Sect. 10.5.3, a non-spherical rigid particle, such as a rod, can move sidewise while settling and so approach the vertical wall. The wall-induced rotation of the particle can lead to a subsequent motion away from the wall. A deformable liquid droplet settling close to a vertical wall will deform into a shape which makes it glide away from the wall"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002237_978-3-319-29357-8_59-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002237_978-3-319-29357-8_59-Figure2-1.png",
        "caption": "Fig. 2 The visualization of relative pose computation between a surface feature and a link of the gripper. In the original work [1] the pose Fi of the link i relative to the feature frame F is used in the PDFs of Contact Model and Query Density (a). We propose that an additional frame Ci, near the contact area, should be introduced during the grasp learning (b) for better approximation of the joint distribution of position and orientation (c)",
        "texts": [
            " The PDF is estimated as: Mi (   p,   q, r ) \u2243 1 Z KO\u2211 j=1 wij3 (   p|   pij, \ud835\udf0ep ) \ud835\udee9 (   q|   qij, \ud835\udf0eq ) 2 ( r|rj, \ud835\udf0ej ) , (2) where Z is the normalizing constant and wij is the weighting function: wij = { exp ( \u2212\ud835\udf06d2ij ) , if dij < \ud835\udeffi, 0, otherwise, (3) where \ud835\udf06 \u2208 \u211d+ is a parameter, dij is the distance from the jth point of the object\u2019s point cloud to the ith gripper link, and \ud835\udeffi is the threshold distance. The weighting function wij causes that only the closest surface features to the gripper link are taken into account. The closer is the feature, the bigger is the weight. The transformation (  pij,   qij) is shown in Fig. 2a. The conditional probability Mi (   p,   q|r ) and marginal probability Mi (r) can easily be computed [1]. Given the Contact Model Mi for the ith link for a specified grasp example and the Object Model O of a novel object, the Query Density PDF Qi can be estimated: Qi(p,   q) \u2243 KQi\u2211 j=1 wij3 (   p|   pQ,ij ) \ud835\udee93 (   q|   qQ,ij ) , (4) where KQi is the number of kernel centers, (   pQ,ij,   qQ,ij ) is the jth kernel center and weights wij are normalized \u2211 j wij = 1. The kernel centers are generated by sampling the object model O (   p,   q, r ) and sampling the Contact Model Mi with condi- tional probability for r: Mi (   p,   q|r ) ",
            " 7) assumes the independence of the position and orientation, and calculates the joint distribution of the pose as the product of two distributions. Such an assumption can be treated as an approximation, but this is generally incorrect. The pose i of the link depends on the surface feature pose j (Sect. 3.3) sampled form the Object Model. In the most cases the distance between the surface feature and the origin of the link frame is large enough, that even small changes in orientation result in relatively large changes in the position of the link w.r.t. the contact point (Fig. 2a). As the sampled pose of the link is dependent on the surface features near the contact point, the translation and the orientation cannot be sampled independently for a given surface feature. We propose a better approximation, that takes into account the position of the contact point (Fig. 2b). During the grasp learning procedure, the mean position of the contact points near the ith link is calculated for each Contact Model:   pi = KO\u2211 j=1 wij Pij\u2215 KO\u2211 j=1 wij, (5) where Pij is the position of the jth closest point on the object\u2019s surface expressed in the frame i, and wij is the weight of the jth contact point. The orientation of the frame  is the same as the orientation of the ith link frame   qi =   qi. The proposed PDF for the Contact Model is estimated as M\u2032 i (   p,   q, r ) \u2243 1 Z KO\u2211 j=1 wij3 (   p|   pij, \ud835\udf0ep ) \ud835\udee9 (   q|   qij, \ud835\udf0eq ) 2 ( r|rj, \ud835\udf0ej ) (6) The difference between Mi and M\u2032 i is that M\u2032 i estimates the probability of finding the ith contact point in the pose i w.r.t. the frame j of the jth local surface feature (Fig. 2c). The ith link pose i can be calculated by combining transformations (   p,   q ) and (   pi,   qi )\u22121 . The Query Density should also be changed to take into account the contact point frame Ci: Q\u2032 i(   p,   q) \u2243 KQi\u2211 j=1 wij3 (   p|   pQ,ij ) \ud835\udee93 (   q|   qQ,ij ) , (7) where (   pQ,ij,   qQ,ij ) is the jth kernel center for the pose of the contact point. The kernel centers are generated by sampling the object model O (   p,   q, r ) and sam- pling the Contact Model M\u2032 i with conditional probability for r: M\u2032 i (   p,   q|r ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002198_1.4033915-Figure23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002198_1.4033915-Figure23-1.png",
        "caption": "Fig. 23 Numbers indicating different regions (description is given in Table 1) for meshing with predefined element sizing option. Red arrow shows the location of applied point load for mesh sensitivity analysis.",
        "texts": [
            " Significant revisions and extensions of the model will be required to enable this next round of simulations. Meshing is one of the most important components of finiteelement simulation. Not only does the mesh have to be of superior quality but also the appropriate element types have to be used. For the current study, various ANSYS Workbench mesh control options were employed to make decision on mesh quality. The model was first divided into nine separate regions to assign different element size options. These nine regions are shown in Fig. 23 and described in Table 1. Following is the strategy based on which element sizing was done. First, the contact region between the hair bulge and socket iris was assigned a contact sizing. This contact area was also geometrically the smallest in the model. Contact sizing in this region was defined as the frictionless contact so that iris can slide freely over the hair budge upon contact. The second set of contact sizing was performed between the hair base and socket. Initially, both the hair and socket were assigned the same material properties, and not much normal stress was observed to be transferred between them",
            " Once the simulation was ran, the structural error was observed to be extremely small [24], further affirming the quality of the mesh. By following the above discretization strategy, four separate models were created by assigning a combination of element sizing options. Element sizing options for these four models are given in Table 1. Discretized models using mesh options \u201cmesh 1\u201d and \u201cmesh 4\u201d is given in Figs. 24(a) and 24(b) for visualization. For all the four models, a point load of 0.5 lN was applied at a location as shown in Fig. 23. Displacements of the loading point as a function of total number of elements for mesh 1, 2, 3, and 4 are plotted in Fig. 25. It is observed that the element sizing option 4 is adequate; however, further finer mesh sizing was employed in the following regions for final analysis: contact region between socket base and hair base with 0.5 lm; socket base and hair base with 0.75 lm; and hair shaft to just past the bulge 0.75 lm. The final discretized mesh is previously shown in Fig. 9. [1] Thurm, U"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001959_sl.2013.3085-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001959_sl.2013.3085-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of inkjet printing of GP-PEDOT:PSS ink on SPCE electrode.",
        "texts": [
            "33 34 Briefly, two graphite rods were plunged in a cell filled with PEDOT:PSS electrolyte and a constant potential of 8 V was applied between graphite rods for 5 hours to generate stable GP-PEDOT:PSS dispersion with a desired graphene concentration. The obtained product was centrifuged at 1200 rpm to separate any large agglomerates and supernatant portion of dispersion was decanted. The GP-PEDOT:PSS solution was employed as an ink for inkjet printing onto SPCE by the commercial material inkjet printer as shown in Figure 1. Four layers of GPPEDOT:PSS were inkjet-printed on the 5 mm\u00d7 25 mm SPCE ver a square ar a of 3 mm\u00d75 mm. The morphology d structure of graphene characterized by transmission electron microscopy (JEOL model JEM-2010), Auger electron spectrometer (AES, EMSL 680), confocal Raman spectroscope (NT-MDT model Ntegra Spectra), scanning electron microscopy (SEM, Hitachi model S-4700), and Fourier transform infrared spectroscopy (FTIR, Perkin Elmer model spectrum spot light-300).33 34 The electrochemical behaviors of inkjet printed GPPEDOT:PSS and unmodified SPCE working electrodes were characterized using the commercial electrochemical work station and home-made electrochemical cell"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000009_icoei.2019.8862541-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000009_icoei.2019.8862541-Figure2-1.png",
        "caption": "Fig. 2. CAD design of proposed robot",
        "texts": [
            " The autonomous agricultural robotic vehicle is controlled by GPS and magnetometer for ploughing, seeding, leveling and the message indication start the irrigation. GPS magnetometer and ultrasonic sensors are used to reach the desired location. Servomotors are used to power the hardware. Wheels powered by DC geared motor are used for movement of the robot. Also, the webcam for live streaming of the ongoing process to the users. Watering is done from the reservoir by the command from the programmed microcontroller [5]. Table I include different kinds of seeds with their depth to be planted [6]. 978-1-5386-9439-8/19/$31.00 \u00a92019 IEEE 1025 seeds (fig. 2). It performs four tasks: \u2022 Intakes the area of the field and forms a grid to plant the seed \u2022 Drills the holes to plant the seed. \u2022 Places the seeds, water them and also provides the fertilizer solution. \u2022 Covers them. Fig.1. Block level diagram of experimental setup The size of the field is transferred to the Arduino using a smartphone and Bluetooth connectivity [4]. Once the grid size and field dimensions are sent, Arduino calculates the grid points to plant the seed. Once reaching the point, the robot drills the field with a blade attached to a 1000 rpm gear motor [9]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure10-1.png",
        "caption": "Fig. 10. Ironing a bulk parallelepiped with a cylinder.",
        "texts": [
            " On the other hand, Table 3 show the results obtained with the consistent normal approach [16,24]. The smoothing quality is noticeably damaged, and the smoothed surface does not properly converge toward the original surface when the number of nodes of the discretized initial shape increases. The relative maximum error is higher, with a value of 0.25 rather than 0.09. A cylinder moves over an elastic bulk parallelepiped (Young\u2019s modulus: E = 2.105 Pa, and Poison\u2019s ratio: m = 0.3) with a horizontal velocity of 1 mm s 1 (Fig. 10). The thickness of the parallelepiped is 5 mm. The mesh size is approximately h = 1 mm. The cylinder radius is 5 mm and its mesh size is approximately h = 1.7 mm. The time step is Dt = 1 s. The simulation was performed with the explicit contact algorithm using different descriptions of tool surface: C0 discretization (facets), analytical description, finer discretization (facets with three time smaller facet size and with edge length of h = 0.5 mm) and smoothed surface. Simulation results first show that smoothing significantly reduces numerical oscillations on forces, which is illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002728_j.intermet.2019.106635-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002728_j.intermet.2019.106635-Figure1-1.png",
        "caption": "Fig. 1. Schematics of processing of tension-interrupted EBSD analysis.",
        "texts": [
            " Intermetallics xxx (xxxx) xxx The 1 A, 3 A, and 5 A specimens were strained in tension at a rate of 1 \ufffd 10 4 s 1 and interrupted at 18% (1st), 36% (2nd), and 54% (3rd) engineering strains, respectively, to examine the deformation microstructure and microtexture using SEM (Zeiss Supra 55, Germany) equipped with EBSD (Oxford Instrument NordlysMax, UK). The equipment was operated at an acceleration voltage of 20 kV with a working distance of 14 mm and an aperture size of 120 \u03bcm. A suitable step size of 0.2\u20132 \u03bcm was selected for scanning according to grain size. The EBSD scanning areas and the corresponding step sizes of each samples are provided in Table 1. The operation procedure of the tensile test and EBSD analysis is shown as Fig. 1. The reason that the equal-spaced strain is decided to characterize the microstructure and microtexture variation after the same amount of deformation and to utilize the total elongation which is at least 50%. All of EBSD data was exported to Channel 5 (Oxford Instrument) software and analyzed for the orientation image maps parallel to the normal direction (ND) and rolling direction (RD) at each strain. The orientation image maps were presented in an inverse pole figures (IPFs) and presented in an inverse pole figure maps (IPF maps), along ND and RD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002033_icelmach.2014.6960180-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002033_icelmach.2014.6960180-Figure9-1.png",
        "caption": "Fig. 9. Time-harmonic FEA simulation result with rotor conductivity set equal to \u03c3=0.021 MS/m.",
        "texts": [
            " However, along the periphery of rotor slots and around the outer rotor circumference, the local conductivity should be set to a much higher value\u2212 approaching MS/m as per hypothesis a) \u2013 due to the local short circuits between adjacent silicon sheets caused by manufacturing defects (punching and cutting). As a consequence, in presence of the mentioned uncertainty, FEA simulations have been repeated according to the two \u201cextreme-case\u201d scenarios a) and b), i.e. assigning the rotor core a uniform conductivity respectively equal to =5 MS/m and =0.021 MS/m. FEA time-harmonic simulation output proved to be highly sensitive to the rotor conductivity value being set in terms of rotor flux distribution, as it can be clearly seen from Fig. 8 and Fig. 9, which refer to the same value of frequency and current. In fact, for the higher rotor conductivity it is clear that a large amount of reaction current arises in the rotor core (as if it were a solid-steel one) causing a sort of \u201cmagnetic shielding\u201d effect which rejects the flux lines towards the outer rotor periphery according to the well-known \u201cskin effect\u201d (Fig. 8). However, it will be seen in the following that the sensitivity of rotor parameters to the rotor axial conductivity values is not so high as could be expected looking at the difference flux distribution (Figs",
            " Since the simulation is a time-harmonic one, a is obviously obtained as a complex number, of which one can determine the module and amplitude: )Re( )Im( arctan a a a (17) 22 )Im()Re( aaa (18) From these values, amplitude and phase of the voltage Vsc\u2019 (Fig. 10) are immediately found as well: asc V ' (19) 2 ' ascV (20) Hence the real and imaginary part of the complex phasor representing Vsc\u2019 are found as: scVscVscV 'cos''Re (21) scVscVscV 'sin''Im (22) This gives an alternative way to compute the parameter Rr_2D as: 3/ 'Re 2_ sc I sc V R Dr (23) It has been checked that the resistance value obtained from (23) is consistent with that obtained through (16). Furthermore, from (22) the total reactance in the equivalent circuit of Fig. 9 is obtained as: 3 Im 2 Isc/ sc V' \u03c3r X D\u03c3s_ XX'eq (24) Since X s_2D is already known from (7), we can compute the rotor leakage reactance as follows: D\u03c3s_ XX' \u03c3r X eq 2 (24) In the 2D FEM simulation it is not possible to evaluate the contribution of the end-winding of the rotor because this cannot be represented in the model. For this reason, the rotor end-winding leakage inductance has been estimated by an analytical formula. First of all the rotor end-winding resistivity at operating temperature is: ))(1( 000 TTf (25) Then the electrical angle between two adjacent bars is computed as: r Z p b 2 (26) where Zr is the number of rotor bars"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000623_s11771-015-2691-7-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000623_s11771-015-2691-7-Figure8-1.png",
        "caption": "Fig. 8 Nyquist plot of Lp(s) for robust stability",
        "texts": [
            " Let p eq( ) ( ) ( )L s G s G s (53) thus, p I I( ) ( ) ( ) ( ) ( )L s L s s L s s   (54) We assume the stability of the nominal closed-loop system, i.e., with \u0394I=0, and the loop transfer function Lp(s) is stable. Then, the stability condition for robust stability of the closed-loop system is tested by using the Nyquist curve. The robust stability condition for ADRIMC system is that if the system is stable, the Nyquist curve of the loop transfer function Lp(s) should not encircle the point \u201c\u22121\u201d. Nyquist curve diagram for Lp(s) is shown in Fig. 8. Then, from Fig. 8, it is observed that the robust stability condition can be described as I ( j ) ( j ) 1 ( j )L L     ,  (55) I ( j ) ( j ) 1 1 ( j ) L L       ,  (56) Assuming that the complementary sensitivity is given by ( ) ( ) 1 ( ) L s T s L s   (57) Substituting Eq. (57) into Eq. (56) yields I ( j ) ( j ) 1T    ,  (58) Equation (58) is equivalent to the following equation: I ( j ) ( j ) 1T     (59) Note that for SISO systems I O( ) ( )s s  and I( ) ( ),T s T s so the condition could equivalently be written in terms of I I( ) ( )s T s or O ( ) ( )s T s "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001798_0954406216631370-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001798_0954406216631370-Figure3-1.png",
        "caption": "Figure 3. Relative motions between unloaded unload gear and output shaft. (a) Pull torque case when _ s 4 _ g ; (b) equal torque case when _ s \u00bc _ g ; (c) drag torque case when _ s 5 _ g .",
        "texts": [
            " Introducing this modulated speed of pinion, the modulated speed government equations of motion are formulated as the following differential equation Ig \u20ac g cos \u00f0 \u00de Xn i\u00bc1 Fi w g, p Rw g g, p \u00bc Td \u00f07\u00de here Ig is the torsional inertia of the gear, is the pressure angle. n is the number of contacted tooth at any given moment. Rw g g, p is instantaneous contact radius of the gear. Fi w g, p is the lubricant reaction force acting on the ith meshing tooth pair. Td is the drag torque between the unloaded gear and output shaft, which is controlled by three conditions as shown in Figure 3. First, the main shaft rotates faster than the unloaded gear. In this case, the drag actually pulls the unloaded gear along and makes it more likely to lose contact with the driving gear Td 5 0\u00f0 \u00de. Second, the main shaft and the unloaded gear rotate with the same speed. There would be no drag torque effect for this rare case Td \u00bc 0\u00f0 \u00de. Third, the last case is when the main shaft runs slower than the unloaded gear Td 4 0\u00f0 \u00de. Throughout this, the transmission process can be either aided or resisted by the drag torque"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000063_wemdcd.2019.8887847-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000063_wemdcd.2019.8887847-Figure6-1.png",
        "caption": "Fig. 6. Configuration of the machine windings and magnets when an angular misalignment of 40% exists (red color: the coils in which the rms value of the EMF is increased, orange color: the coils in which the rms value of the EMF is decreased, in a cycle: the coils in which the rms value of the EMF remains approximately constant): a) in y-axis direction, b) in zaxis direction",
        "texts": [
            " Likewise, the rest coils EMF amplitudes present increment or decrement related to their position, but as they do not give an extra information are not presented. From the table I it can be observed that the EMF rms value of coils A4, B4, B3, B4, C3 and C4 reduces, because these coils located in the position of increased airgap length. The EMF rms value of coils A2, B1, B2, C1 and C2 increases because these coils located in the position of reduced airgap length, while the EMF rms value of coils A1 and A3 remains almost constant because the airgap length remains constant due to angular misalignment. In Fig. 6 the coils with the reduced EMF are presented with orange color, the coils with increased EMF are presented with red color, while the coils with EMF roughly constant are presented in a cycle. The FFT spectra of the EMF of the coil of phase A of the machine are depicted on Fig. 7. The fault does not create new harmonics in the spectra. The existing harmonics in both healthy and faulty case are of orders 3th, 5th and 7th due to the harmonic components that exist on the magnetic field due to the winding configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001776_j.ifacol.2015.06.362-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001776_j.ifacol.2015.06.362-Figure1-1.png",
        "caption": "Fig. 1. Formate machine tool and workpiece setup.",
        "texts": [
            " Section 3 shows how to derive the working rake and relief angles which play crucial roles in tool wear. In section 4, the tool wear is characterized. The new blade modeling method is presented in section 5. Finally, in section 2. KINEMATIC STRUCTURE OF FORMATE (NONGENERATED) FACE-HOBBING AND WORKPIECE SETUP In Formate process, cradle does not have any rotation. The cutting system and workpiece rotate proportionally and cutter has a feed movement along its axis, so two rotations and one translation. Structure of the machine tool is illustrated in Fig. 1. Coordinate system Sh is rigidly connected to the head cutter and rotates with it. The head cutter rotation is measured (\u03b8) in coordinate system S1. Z1 (z axis of S1) is common with Zh and plane XY1 (xy plane of S1) is coincident with XYh. At the begining of the machining process, the head cutter has an offset to back (BO) of the machine plane (XYm) and during machining, cutter head is fed toward the machine plane (XYm) by an axial translational movement (Feed in Fig.1). Z1 is perpendicular to the plane XYm. The workpiece coordinate system (Sw) is rigidly attached to the workpiece and rotates with the workpiece. Machine settings are adjusted in Sm (machine coordinate system). For Formate process, four machining settings are required (Fig.1), horizontal setting H, vertical setting V, machine root angle \u03bbm and machine center to back \u0394Xp. In addition, the workpiece rotates proportionally with head cutter with a ratio as Rb= Nb/Ng, where Nb is number of blade groups and Ng is number of gear teeth. Workepiece rotation is measured in coordinate system S2. Z2 is common with Zw and XY2 is coincident with XYm. Because the workpiece and the head cutter have relative motion with respect to each other, in order to know the effect of blades when they remove material from workpiece, the representation of the blade cutting edge should be derived in the workpiece coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure13-1.png",
        "caption": "Figure 13. The extreme case of moving material away from the neutral surface to carry bending loads more efficiently. (a) Material moved entirely to top and bottom face. (b) When loaded, upper and lower plates pinch together, bringing them closer to the neutral surface and increasing stress. (c) To keep the upper and lower surfaces a constant distance away from the neutral surface, engineers add a vertical \u2018web\u2019 to parallel bars to produce the standard I-beam.",
        "texts": [
            " In fact, for resisting bending, torsion, and Euler buckling, a hollow cylinder can be over an order of magnitude stiffer that a solid one of the same cross sectional area (and thus, same mass). The farther the material can be placed from the neutral axis, the stiffer the strut will be, except that if the wall gets too thin local buckling can become a danger. Consider a beam that is only subject to bending in one plane, such as a horizontal cantilever supporting a weight at one end. The most efficient arrangement of material would be to have it concentrated at the top surface\u2014to resist tension\u2014and bottom surface\u2014to resist compression\u2014which amounts to two parallel, flat bars (figure 13). When loaded, however, the upper and lower bars would tend to pinch together, reducing their I and hence their bending stiffness. This is why engineers use I-beams: the vertical \u2018web\u2019 keeps the top and bottom surfaces spread apart to maintain a high I. Many struts must resist bending from arbitrary directions, so a hollow, circular tube is the most efficient arrangement of material. Moreover, a hollow tube also turns out to be the most efficient shape for resisting torsion, due to the conceptual similarity of J and I and the need to maximize J"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000747_acc.2013.6580230-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000747_acc.2013.6580230-Figure1-1.png",
        "caption": "Fig. 1. Interconnected pendulums.",
        "texts": [
            " To simplify the calculation, we establish the following theorem which requires only one LMI to be feasible, as follows       Z\u0304 Y (\u03bb\u0304Q)1/2 Y Y (\u03bb\u0304Q)1/2Y \u2212I 0 0 Y 0 \u2212 1 \u03c0 I 0 Y 0 0 \u2212 1 (N\u22121)\u03b8 I       < 0, (25) where p2 = max i (pi) 2, q2 = max i (q2i ), \u03bb = min i (\u03bbi), \u03bb\u0304 = max i (\u03bbi), and Z\u0304 = AY + Y A\u2032 \u2212 \u03bb2 \u03bb\u03042 B1R \u22121B\u20321 + [p2 \u03c0 + q2 \u03b8 ] B2B \u2032 2. Theorem 2: Given R = R\u2032 > 0 and Q > 0, suppose the LMI (25) in variables Y = Y \u2032 > 0, \u03c0\u22121 > 0 and \u03b8\u22121 > 0 is feasible. Then the control protocol (5) with K = \u2212 \u03bb \u03bb\u03042 R\u22121B\u20321Y \u22121 solves Problem 1. Furthermore, this protocol guarantees the following performance bound sup \u039e0 J (u) \u2264 N(\u03c0+ \u03b8(N \u2212 1))d+ N \u2211 i=1 e\u2032i(0)Y \u22121ei(0). (26) To illustrate the proposed method, consider a system consisting of three identical pendulums coupled by two springs. Each pendulum is subject to an input as shown in Fig. 1. The dynamic of the coupled system is governed by the following equations ml2\u03b1\u03081 =\u2212 ka2(\u03b11 \u2212 \u03b12)\u2212mgl\u03b11 \u2212 u1, ml2\u03b1\u03082 =\u2212 ka2(\u03b12 \u2212 \u03b13)\u2212 ka2(\u03b12 \u2212 \u03b11) (27) \u2212mgl\u03b12 \u2212 u2, ml2\u03b1\u03083 =\u2212 ka2(\u03b13 \u2212 \u03b12)\u2212mgl\u03b13 \u2212 u3, where l is the length of the pendulum, a is the position of the spring, g is the gravitational acceleration constant, m is the mass of each pendulum, and k is the spring constant. In addition to the three pendulums, consider the leader pendulum which is identical to those given. Its dynamics are described by the equation ml2\u03b1\u03080 = \u2212mgl\u03b10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003994_978-3-319-10924-4_2-Figure2.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003994_978-3-319-10924-4_2-Figure2.1-1.png",
        "caption": "Fig. 2.1 This figure shows the original scenario that motivated the need for self-reconfigurable robots (Courtesy of Fukuda, \u00a9 1988 IEEE)",
        "texts": [
            " The idea was that if multiple robots could automatically form physical bonds between each other the combined robot collective could adapt its shape and functionality in response to the environment and tasks. The basic scenario K. Stoy (B) IT University of Copenhagen, Rued Langgaards Vej 7, 2300 Copenhagen S, Denmark e-mail: ksty@itu.dk \u00a9 Springer International Publishing Switzerland 2015 G.Ch. Sirakoulis and A. Adamatzky (eds.), Robots and Lattice Automata, Emergence, Complexity and Computation 13, DOI 10.1007/978-3-319-10924-4_2 33 that motivated self-reconfigurable robotic research given by Fukuda et al. [8], shown in Fig. 2.1, was that individual robots couldmove into a storage tank through a narrow passage and once inside they could assemble for the purpose of cleaning the storage tank. This vision of self-reconfigurable robots was and is still today attractive. However, the scientific challenges involved in realising this vision are significant. One aspect is the mechatronic realisation of self-reconfigurable robots and another central to the topic of this chapter is the question of their control. Inmany self-reconfigurable robotsmodules are organised in a lattice structure like atoms in a crystal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000803_j.actaastro.2014.04.009-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000803_j.actaastro.2014.04.009-Figure10-1.png",
        "caption": "Fig. 10. Experimental apparatus of the planar dual-link robot equipped with a flexible appendage.",
        "texts": [
            " In this study, in order to validate the effectiveness of the proposed method, which takes into consideration the dependency of the frequency on the link angles, two cases are investigated. In the first case, we focus on the first modal vibration at only the link angles \u03d51 \u00bc \u03d52 \u00bc 901 and apply the inputshaping technique to this frequency. In the second case, the time-delay interval of the input shaper is adjusted in accordance with the modal frequency estimated from the current link angles during the control process. Fig. 10 shows the experimental setup of a planar duallink space robot equipped with a flexible appendage that supports a mass at its free end. This robot is equipped with a magnetometer to detect the attitude of the main body, two stepper motors to drive each link angle, and two encoders to detect each link angle. Note that the operational angle of each link is restricted within 71101 due to structural limitations. A large glass board, referred to as a flight-bed, is placed horizontally. In order to simulate microgravity, air bearings are placed under the end-effector, the link joint, and the main body of the robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.2-1.png",
        "caption": "Figure 3.2 Torque produced by coil in magnetic field",
        "texts": [
            " The work of Jean Baptiste Biot (1774\u20131862), and Felix Savart (1791\u20131841), formulated the relationship between current in a conductor and the magnetic field surrounding it, which led to Hedrick Antoon Lorentz (1853\u20131928) deriving the formula for the force on a current-carrying conductor in a magnetic field, as illustrated in Figure 3.1. F = BlI (3.1) where B is the magnetic field strength, l is the length of the current-carrying conductor, I is the current magnitude, and F is the Lorentz force. If a flat coil of wire with radius r is placed in a magnetic field and mounted on an axis, as shown in Figure 3.2, then current flowing through the coil, with the polarities shown, will result in Lorentz forces developing on the coil wires, leading to creation of torque and rotation of the coil, where: T = BlIr sin \u03b8 (3.2) As the coil rotates counter clockwise (CCW), the torque will vary in a sinusoidal fashion from a maximum (\u03b8 = 90\u25e6) to zero (\u03b8 = 0\u25e6) at which point motion would stop. However, if a second coil is mounted at an angle to the first (dotted in Figure 3.2), then it would create additional torque to maintain motion and rotate the first coil such that \u03b8 will increase beyond 0 and would again develop torque if its current were reversed in order to maintain CCW rotation. The reversal of the current in the rotating coils is defined as commutation. System Components 37 A typical motor will have numerous coils mounted on the rotor (the armature) connected to a series of copper segments, the commutator, mounted on one end of the rotor. Current flows through the coils via these segments that in turn are contacted by stationary soft carbon devices (the brushes)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001390_cicare.2014.7007838-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001390_cicare.2014.7007838-Figure4-1.png",
        "caption": "Figure 4: An illustration of door sensor and its implementation components",
        "texts": [
            " Similarly to the mats, the pressure sensors can be embedded in chairs, sofas, etc. to determine the patient\u2019s position and trace his/her activity. Door sensors (DS) are employed to alert a caregiver when a monitored person exits his room and/or opens an entrance door. Unlike existing DS that just sense the door opening and closing, our sensor detects who opens/closes the door and alerts the caregiver when it is done by person with cognitive impairment (PCI). The device combines a magnetic sensor with a radio-frequency identifier of the patient, as shown in Fig.4 (a). To identify the PD, we embed a RFID passive tag in his/her slippers, as shown in Fig.4 (b) bottom-left, and place antenna of active RF-reader (Fig.4 (b), right image) under the carpet mat in front of a door, similarly as in [20]. However, unlike [20], our RFID readers are wireless. Both the magnetic sensor and the RF-reader are connected to the Xbee transmitters, which signal the server whether the door is open or close and whether the person at the door is the PD or not, respectively. Based on this pair of signals, the server (see Fig.4,a) assesses if the door is opened by the PD. The tests confirmed that both tag and the RFID antenna are not perceptible as they walked. Also, using slippers as RFID-tag is an acceptable solution, since elderly people usually wear slippers at home all the time. The passive RFID tags do not require charging batteries minimizing the maintenance tasks. The bed sensor (BS) detects the presence of a person in bed. It consists of an air bag placed between the mattress and the base in the patient\u2019s bed, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001194_s00366-013-0350-x-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001194_s00366-013-0350-x-Figure3-1.png",
        "caption": "Fig. 3 The diagram of the first mesh of the gear teeth",
        "texts": [
            "4, the number of teeth is 11, the pressure angle is 20 . The manufacturing material of the gear is C45, the elastic modulus of it is 200 GPa, the Poisson ratio of it is 0.3, and the density of it is 7,859 kg/ mm3. Based on the blind source separation with the particle swarm optimization algorithm, the natural frequencies are obtained. The natural frequencies of intact and cracked gear pump gear are listed in Table 1. The part of the gear pump gear tooth is divided into 11 zones, and the wavelet finite element division is shown in Fig. 3. The consistent meshing technique is used to ensure the results credible, and the mesh size is reduced in the vicinity of the crack for every cycle, and the curved boundary is meshed by curved finite elements and straight line elements. When the expression (20) is substituted into expression (25) and the effect coefficient matrix H~can get. According to the data in Table 1, the following expression can be expressed as follows: Dx x6 1 \u00bc 2H~6 11D~11 1 \u00f032\u00de The damage coefficient vector [d1, d2, d3, , d11]T can be obtained through solving the expression (32), which is listed as follows: D~ \u00bc \u00bdd1; d2; d3; d4; d5; d6; d7; d8; d9; d10; d11 T \u00bc \u00bd0:0005; 0:0027; 0:0024; 0:0017; 0; 0; 0; 0; 0:0023; 0:0002; 0:0003 T Where d3 and d9 are \\0, and they are set as 0, and the expression (32) is solved again, and the new damage coefficients are obtained as follows: D~ \u00bc \u00bdd1; d2; d4; d5; d10; d11 T \u00bc \u00bd 0:00036; 0:0027; 0:0014; 0:0032; 0:0018; 0:00021 T Where d1, d2, d5 and d10 are\\0, and they are set as 0, and the following results are as follows: D~ \u00bc \u00bdd4; d10 T \u00bc \u00bd0:0012; 0:0003 T According to the final results, the elements of the damage matrix are all positive value, therefore, the zonesP e4 and P e10 can be predicted as cracked elements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000200_iecon.2019.8927009-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000200_iecon.2019.8927009-Figure1-1.png",
        "caption": "Fig. 1. Pepper coordinate system using ALMotion module.",
        "texts": [
            " This means that when the method is running no other method can run simultaneously, implicating that the point definition would have to be done separately and after exploration. To deal with this issue, the objective passed by being able to run each module at the time in order to obey the blocking calls and run the experiments smoothly. While constructing the map, the robot creates a coordinate system relative to the position where Pepper starts the mapping process. When using the ALMotion module Pepper uses a coordinate system that consists in 3 coordinates [x, y, j] being theta the rotation around z-axis in radians Fig. 1. However, when using the Navigation functions, Pepper does not use j, 5265 Authorized licensed use limited to: University of Exeter. Downloaded on June 21,2020 at 07:29:36 UTC from IEEE Xplore. Restrictions apply. but instead it only uses xy coordinates relative to the map. Pepper can also choose its own path and speed while moving around. Like exploration, navigation also represents a blocking call. After the mapping process is completed, since the navigation functionality only uses two coordinates and the orientation cannot be controlled, we felt the need of marking various points on the map in order to make good use of the navigation and obstacle avoiding methods"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003590_ecce44975.2020.9235871-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003590_ecce44975.2020.9235871-Figure13-1.png",
        "caption": "Fig. 13. The CHTC of the recirculating hollow-shaft at a speed of 3000rpm and a flow rate of 5L/min. (a) original structure, (b) optimized structure for reducing inner diameter of tube.",
        "texts": [
            " However, the CHTC in the bottom area is about 500 W/m2/K when the bottom is tapered, which is twice that before optimization. Although changing the bottom to a cone-shape can improve the CHTC of the bottom area, the overall CHTC does not increase much. A simple optimization method is to gradually reduce the inner diameter of the bottom stationary cooling tube, so that it enhances the disorder of the bottom fluid, thereby improving the value of the overall average CHTC. The detailed structure is depicted in Fig. 12. Fig. 13 shows that the maximum value of the CHTC of the structure after optimization is 1.57 times that before optimization, and the average CHTC is increased by 16%. However, the friction loss of the wall surface increased by 78%. In addition, the pressure drop of the system is greatly increased due to the reduction of the inner diameter of tube. Obviously, from the overall consideration, reducing the inner diameter of tube is not a good optimization method. 3515 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003568_10.0001659-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003568_10.0001659-Figure12-1.png",
        "caption": "Fig. 12. Illustration of the ball launch machine. Varying the spin rate at constant launch speed is achieved by the condition jxAj \u00fe jxBj \u00bc constant.",
        "texts": [
            " Humans are usually not capable to just vary the spin rate while keeping the launch angle and the speed constant. In the ball launch machine, two spinning wheels accelerate the volleyball quickly to the launch speed. As the two wheels can be spun at different spin rates, spin can be imparted to the volleyball. The launch angle is kept constant by fixing the orientation of the machine. The launch speed is kept constant by assuring that the sum of the magnitudes of the spin rates of the two wheels is constant during the experiment (Fig. 12). Figure 13 shows a comparison of the trajectories of volleyballs with various rates of topspin and backspin which are described by the number of rotations per seconds (rps). We also plot a vacuum trajectory for reference. The larger the topspin, the larger the downward Magnus force and the faster the ball dips below the trajectory of the ball without spin. The larger the backspin, the stronger the upward Magnus force and the longer the range of the ball becomes. For backspin, the Magnus force becomes a lift force opposing gravity, as for an aircraft wing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000741_jfm.2014.45-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000741_jfm.2014.45-Figure6-1.png",
        "caption": "FIGURE 6. (Colour online) The contours of the spanwise component of the virtual velocity (\u2207\u03d5T) caused by the virtual moving right-rear goose along the thrust direction (z), where the curves approximate to circle arcs when \u03b8T \u2208 [\u03c0/6,\u03c0/3].",
        "texts": [
            " First, we consider the thrust on the right rear goose moving virtually along the thrust (forward) direction. Here T4 is employed to predict the contribution of the leading goose\u2019s wake to the thrust. The virtual velocity potential excited by the forward virtual moving \u2018goose\u2019 (right-rear in figure 5) approximates to \u03d5T =\u2212u\u22170 a3 2r2 cos \u03b8T, (4.4) where a is the radius of the sphere, r is the radial distance from the centre of the goose\u2019s body to the point we considered and \u03b8T is the azimuthal angle between the zenith (forward) direction and the vector r (see figure 6). Then, the magnitude of the spanwise component of the virtual velocity is v\u2217 =\u22133u\u22170 2 (a r )3 sin \u03b8T cos \u03b8T, (4.5) which is located in the horizontal plane. In the above formula, \u2018\u2212\u2019 and \u2018+\u2019 indicate the left and right front regions of the goose, respectively. As a preliminary study, substitute (4.5) into (4.3), and set the distance, d, between the right-rear goose and the near wake of the leading one to be constant, then we obtain T4 \u221d \u0393 2 4\u03c0b 3u\u22170 4 (a d )3 sin(2\u03b8T), (4.6) which is positive when \u03b8T \u2208 (0, \u03c0/2)",
            "7) According to the above equation, the optimal azimuthal angle is \u03b8T =\u03c0/4, where the goose obtains the maximum drag reduction, Dr0. Furthermore, near \u03c0/4, the value of the drag reduction does not change a lot. Specifically, when \u03c0/66 \u03b8T 6\u03c0/3, the drag reduction is more than 86 % of the maximum drag reduction Dr0, i.e. \u221a 3/2 6 sin(2\u03b8T)6 1, when \u03b8T \u2208 [\u03c0/6,\u03c0/3] (4.8) Thus, it can be concluded that the right-rear goose can receive the induced thrust (the drag reduction) when the front goose flies at the 10 to 11 o\u2019clock position (see figure 6). According to the geometry relation, the angle of the V formation flight of geese, \u03b1 = 2\u03b8T , is \u03b1 \u2208 [\u03c0/3, 2\u03c0/3]. (4.9) On the other hand, there may be negative effect for the leading goose if the trailing goose receives a positive effect. Certainly, by using the same analysis method, the virtual power T4 is negative for the leading goose caused by the wake of the rightrear or the left-rear in the V formation. It reveals that the leading goose experiences larger drag than it flies solely. According to (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure6-1.png",
        "caption": "Fig. 6: Kinematic chain selection",
        "texts": [
            " In this paper the downward or ship welding attitude, which can inhibit the flowing of iron melt well, are adopted to ensure the welding quality. These two ideal welding attitudes are drawn in Fig. 5. Two basic assumptions are made here. The relative pose transformation between the tool and the workpiece is not considered here. Instead the transformations between tools of multiple robots are taken into account via cooperation constraints. Moreover, considering tools are always fixed to the end of flange, the kinematic chain of multi-robot system composed by a master robot and several slave robots can be simplified as Fig. 6 shows. For the master robot, the kinematic chain from the master robot base to the master robot tool is considered. For slave robots, the kinematic chain is chosen from the master robot base to the slave robot base, and to the slave robot tool, and finally to the master robot tool. In Fig. 6, the red dashed line represents the master robot kinematic chain and yellow dashed line represents a slave robot kinematic chain. They constitute a closed kinematic chain. {wp} stands for the workpiece coordinate system. {m} and {s} are the master and slave robot base coordinate system respectively. {m, tcp} and {s,tcp} are the coordinate system of master and slave robot tools. According to the operation requirement of cooperated multi-robot system, the closed kinematic chain of master-slave robot can be expressed as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003738_j.promfg.2020.11.026-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003738_j.promfg.2020.11.026-Figure3-1.png",
        "caption": "Fig. 3. Results of initial FEM simulation of distortion compensation via locally adjusted heat transfer coefficients: (a), homogeneous HTC over full surface, (b) HTC reduced by a factor of ten on lower, (c) on upper surface. Distortion and maximum principal stress after cooling depicted.",
        "texts": [
            " The aim is to reduce latency to such a degree that predictions of distortion are available in real time. This would allow compensating measures to be adapted to each individual component, and thus to cover not only systematic process deviations, but also those based on parameter fluctuations [3]. Feasibility of the fundamental approach as described in the preceding section was evaluated based on 2D FEM simulations of a geometrically simple reference part specifically chosen to exhibit distortion under conditions of homogeneous cooling, i.e. identical HTC values over the full surface area. Fig. 3 and 4 depict the results in terms of distortion and stress state (Fig. 3) as well as equivalent plastic strain (Fig. 4) after cooling to room temperature. The simulation is based on temperature-dependent stress-strain data for the alloy GDAlSi10Mg. Simulations start at a uniform temperature of 500\u00b0C, roughly matching typical solution heat treatment temperatures, throughout the part. Temperature dependence of thermal conductivity and heat capacity were modeled using two interpolation values sourced from supplier specifications, while HTCs were assumed constant over the temperature range covered. Fig. 3a and 4a show the effect of homogeneous cooling at an HTC of 120 W/(m2K), which matches typical values for forced convection cooling with air [17]. The deformation observed meets expectations in as far as the two arms of the structure move inwards due to the belated cooling and shrinkage of the material aggregation between them. In the case of Fig. 3b and 4b, the HTC on the upper surface was maintained, while it was reduced by a factor of ten for the lower. As this causes a further delay in cooling of the aforementioned material agglomeration, the distortion observed in the initial case is increased. Fig. 3c and 4c reflect a reversal of this situation, and consequently also a reversal in the orientation of distortion: The effect of homogeneous heat extraction is not only nullified, but in fact overcompensated, as the structure\u2019s two arms bend outwards in this case. The color coding in Fig. 3 illustrates the maximum principal stress after cooling, i.e. the residual stress in the component. While case (a) results in similar compressive stress levels at top and bottom surfaces, inhomogeneous cooling (b, c) shifts the dominant compressive stresses to the surface associated with the higher HTCs, while the area fraction of compressive stresses is generally reduced. 148 Alireza Ebrahimi et al. / Procedia Manufacturing 52 (2020) 144\u2013149 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 5 Figure 4 shows the accumulated deformation, expressed as equivalent plastic strain, at the end of the cooling step",
            " / Procedia Manufacturing 52 (2020) 144\u2013149 149 6 A. Ebrahimi et al. / Procedia Manufacturing 00 (2019) 000\u2013000 system [19]. At the same time, as a capability and as part of an inverse process model, it is a prerequisite for the planned translation of locally required HTCs to spraying system settings and will be addressed by means of combined experimental and CFD simulation based studies. Comparison of the HTC values in Fig. 6 with the respective assumptions in section 4 (see specifically data assumed in the simulations behind Fig. 3 and 4) show that there is ample space for local variation of cooling conditions. Controlling distortion is a major challenge in high pressure die casting. Furthermore, with the present move of the HPDC industry into the realm of structural automotive components, its significance is bound to increase. Conventional approaches to nevertheless achieve dimensional accuracy by means of reworking are costly, while embodying previous, process simulation-based knowledge on the dimensional deviations to be expected in hardware, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure12-1.png",
        "caption": "Figure 12. Velocity contours for modified body geometry along the Y-axis.",
        "texts": [],
        "surrounding_texts": [
            "Table 1 and the graph (Figure 14) show the calculations of the maximum speed of the velocity fields at fixed drum and beater revolutions 640 rpm and 2100 rpm, respectively, and different values of the drum diameter. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 At drum diameters less than 550 mm, the maximum speed values are preserved, since the maximum flow rates in these cases are already set by the beater and remain unchanged. Figure 14. Velocity field contours for drum diameter 350 mm. Figure 15. Velocity field contours for a drum diameter of 450 mm. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0003756_wre51543.2020.9307007-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003756_wre51543.2020.9307007-Figure5-1.png",
        "caption": "Fig. 5: Overview of the positioning of sensors and components of the simulated robot.",
        "texts": [
            " This robot structure is written in the Unified Robot Description Format (URDF) [17], an XML specification file that contains the complete physical information of link, joints, properties, actuators, and sensors. It allows the use of computational processes such as Computer-Aided Engineering (CAE) and Computer-Aided Design (CAD) to facilitate the import of objects for simulation. The CoppeliaSim simulator already offers several sensors and equipment used in robotics natively. However, for a closer simulation of the real robot, some sensors were adapted from these structures already available. The representation of the virtual robot with all sensors is illustrated in Figure 5. Cameras: For the cameras simulation, the Vision Sensor structure of CoppeliaSim was used, which provides image rendering by frame and transmitting depth data. From there, the images and data for each camera are published in ROS. To view these images in the developed graphic interface, compatibility with video streaming protocols such as Real Time Streaming Protocol (RTSP) is required. Thus, the webvideo-server package, developed for streaming in HTTP and RTSP protocol from ROS image, was used. RealSense D435i: This sensor was developed from adaptations in the model of Kinect, RGBD camera already available in the simulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure2-1.png",
        "caption": "Fig. 2. (a) Appearance diagram of a general homopolar-type MB. (b) Bias flux of a general homopolar-type MB.",
        "texts": [
            " Thus, the hetelopolar-type MB is often employed in a high-output motor rotating at high speeds. However, in the case of the heteropolar-type MB, a large iron loss of a rotor is caused by an alternating magnetic field on a rotor core. Consequently, the rotor gets heated drastically by the large rotor iron loss. In an MB, it is difficult to cool a rotor owing to the mechanical noncontact. Therefore, in terms of cooling, the heteropolartype MB has a problem of large iron loss, causing heating of the rotor. On the other hand, the general homopolar-type MB structure, shown in Fig. 2(a), has low rotor iron loss, Manuscript received March 20, 2015; revised July 4, 2015; accepted July 6, 2015. Date of publication July 14, 2015; date of current version October 22, 2015. Corresponding author: T. Matsuzaki (e-mail: matsuzaki@sfc.ssi.ist.hokudai.ac.jp). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2015.2456338 Fig. 1. Heteropolar-type MB and its bias flux. because the magnetic field on a rotor core becomes dc due to the magnetic flux flowing radially from a center of a rotor shaft, as shown in Fig. 2(b). Therefore, in order to decrease the rotor iron loss and make the cooling of the rotor easy, the general homopolar-type MB is employed. However, in order to stably generate sufficient suspension force for suspending a shaft of a high-output motor, the general homopolar-type MB requires a large magnet and arch-shaped windings, as shown in Fig. 2(a). It is noted that the large magnet is required for generating the bias flux as the dc magnetic field and the arch-shaped windings is needed to achieve the high fill factor. Consequently, the general homopolar-type MB has a complicated structure and becomes costly. In order to realize the MB structure with both lower iron loss and lower cost than the general homopolar-type MB, our research group focuses on the homopolar-type MB structure using four C-shaped cores, as shown in Fig. 3 [2]\u2013[4]. The homopolar-type MB structure with the four C-shaped cores has potential for lower rotor iron loss than the general homopolar-type, because the magnetic field on a rotor core is dc, and the width between the magnetic poles can be smaller than that of the general homopolar-type, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002621_b978-0-12-800806-5.00014-7-Figure14.1-1.png",
        "caption": "FIGURE 14.1 A two\u2013member frame: model and load.",
        "texts": [
            " To summarize, a general formulation for the design problem involving explicit and implicit functions of design variables is defined as: find an n-dimensional vector x of design variables to minimize a cost function f(x) satisfying the implicit design constraints of Eq. (14.2), with U satisfying the system of Eq. (14.1). Note that equality constraints, if present, can be routinely included in the formulation as in the previous chapters. We illustrate the procedure of problem formulation in Example 14.1. K(x)U=F(x) gix, U\u22640 II. NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION EXAMPLE 14.1 DESIGN OF A TWO\u2013MEMBER FRAME Consider the design of a two\u2013member frame subjected to out-of-plane loads, as shown in Fig. 14.1. Such frames are encountered in numerous automotive, aerospace, mechanical, and structural engineering applications. We want to formulate the problem of minimizing the material volume of the frame subject to stress and size limitations (bartel, 1969). Since the optimum structure will be symmetric, the two members of the frame are identical. Also, it has been determined that hollow rectangular sections will be used as members with three design variables defined as: d = width of the member, in. h = height of the member, in",
            " With this criterion, f(x)=2L2dt+2ht\u22124t2 II. NUMErICAl METhODS FOr CONTINUOUS VArIAblE OpTIMIzATION the effective stress \u03c3e is given as \u03c3 \u03c4+ 32 2 and the stress constraint is written in a normalized form as: \u03c3 \u03c3 \u03c4( )+ \u2212 \u2264 1 3 1.0 0 a 2 2 2 (14.4) where \u03c3a is the allowable design stress. The stresses are calculated from the member-end moments and torques, which are calculated using the finite\u2013element procedure. The three generalized nodal displacements (deflections and rotations) for the finite\u2013element model shown in Fig. 14.1 are defined as: U1 = vertical displacement at node 2 U2 = rotation about line 3\u20132 U3 = rotation about line 1\u20132 Using these, the equilibrium equation (Eq. (14.1)) for the finite\u2013element model that determines the displacements U1, U2, and U3, is given as (for details of the procedure to obtain the equation using individual member equilibrium equations, refer to texts by haug and Arora, 1979; Chandrupatla and belegundu, 2009; bhatti, 2005): \u2212 \u2212 +    +                              =         EI L L L L L GJ EI L L L GJ EI L U U U P 24 6 6 6 4 0 6 0 4 0 0 3 2 2 2 2 1 2 3 (14",
            " The user-supplied subprograms can be quite simple for problems having explicit functions and complex for problems having implicit functions. External programs may be called upon to generate the function values and their gradients needed by the optimization program. This has been done, and several complex problems have been solved and reported in the literature (lim and Arora, 1986; Thanedar et al., 1986, 1990; Tseng and Arora, 1988, 1989). 14.5 OPTIMUM DESIGN: TWO\u2013MEMBER FRAME WITH OUT-OF-PLANE LOADS Fig. 14.1 shows the two\u2013member frame subjected to out-of-plane loads. The members of the frame are subjected to torsional, bending, and shearing loads. The objective of this problem is to design a structure having minimum volume without material failure due to applied loads. The problem was formulated in Section 14.1.2 using the finite\u2013element approach. In defining the stress constraint, the von Mises yield criterion is used and the shear stress due to transverse load is neglected. The formulation and equations given in Sections 14"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002518_15325008.2016.1236852-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002518_15325008.2016.1236852-Figure2-1.png",
        "caption": "FIGURE 2. Radial force magnitude variation in each mechanical position produced only by torque windings currents. (a) Rotor shaft is no eccentricity. (b) Rotor shaft eccentricity displacement of 0.05 mm in both directions.",
        "texts": [
            " If the external force or other disturbance factors cause rotor to move away from the equilibrium position, the currents of the X - or Y -axis suspension windings have to be properly controlled to force the rotor back to the center of the stator. The rotor shaft eccentricity causes the phenomenon of the non-uniform air gaps between the rotor and stator poles. Furthermore, rotor net radial suspending force, which is produced only by torque windings currents, has been no longer balanced. With FEM analysis, Figure 2 shows varied magnitude curves of rotor radial net force when rotor rotates a round clockwise. The BSRM stator/rotor pole number is 12/8, all torque windings ampere-turn is 72 A when energized, the suspension windings ampere-turn is 0 A, and the air-gap length is 0.5 mm in the FEM simulation model. In Figure 2(a), the rotor shaft is located in the central position, it has no eccentricity, the radial force is increased with the increase of the overlapped area of stator and rotor tooth poles, it is in the same way for each of the conducting phase. Figure 2(b) shows the result when there are both 0.05-mm displacements in the positive \u03b1-direction and in the positive \u03b2-direction. In Figure 2(a), the radial force has the same variation pattern for each energized phase. The values of the radial force are different with the variation of the conduction angle, which is the overlapped area of stator and rotor tooth poles. In Figure 2(a), 0\u25e6 \u223c 15\u25e6, the phase-B is energized. For example, in the 0\u25e6 stator position, that is the overlapped area of phase-B stator and rotor tooth poles is zero, the rotor radial force is minimum 0 N. In the 15\u25e6 stator position, that is the overlapped area of phase-B stator and rotor tooth poles is maximum, the rotor radial force also has maximum value 25 N. But in Figure 2(b) when rotor shaft is located in the eccentricity position, the radial forces situation has been no longer the same for the different energized phases. As shown in Figure 2, the rotor shaft eccentricity has definitely caused a great change of the rotor radial force even though the suspension windings have not been energized, and there are different changes according to the energized phase winding. Therefore, the rotor eccentricity has to be considered to build the radial forces analytical model of BSRM. In the following analysis, magnetic equivalent circuit method is employed, and coupling between the torque windings is ignored. The neglect of coupling is possible, because for BSRM there is only one phase torque winding being energized at any moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003476_0954408920948683-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003476_0954408920948683-Figure3-1.png",
        "caption": "Figure 3. Kinematics nomenclature of continuum structure.",
        "texts": [
            " The parameters defined as below: ri Projection of the driving wire to the YAOAZA plane L Length of continuum structure (L\u00bc 280mm) Li Length of the driving wire DLi Variation of the driving wire hi Distance from Li to ZA-axis qi The radius of curvature of Li h Bending angle of continuum structure u Rotation angle of continuum structure a Rotation angle of turntable hT Displacement of vertical sliding table qT Displacement of horizontal sliding table The continuum structure is driven by three wires, which are fixed to I, II and III position of the top tendon guide, and cross through the hole of other guides, as shown in Figure 2. In order to establish the relationship between the variation of the driving wires and the position of the end of continuum structure, the parametric kinematics analysis was performed by using constant curvature mode,27,28 as shown in Figure 3. Base coordinate system (A) x\u0302A; y\u0302A; z\u0302A is the original coordinate system. Rotation coordinate system (B) x\u0302B; y\u0302B; z\u0302B is obtained by rotating of the A coordinate system around the ZA-axis for u degrees. System (C) x\u0302C; y\u0302C; z\u0302C is the free end coordinate system, the angle between XC-axis and XA-axis is h. Free end rotation coordinate system (D) x\u0302D; y\u0302D; z\u0302D is obtained by rotating of the C coordinate system around the ZC-axis for u degrees, and the XD axis is parallel to the XA-axis. The relationship between the coordinate of the end point of the continuum structure in three-dimensional space and its rotation angle and bending angle can be deduced by using the trigonometric function xD yD zD 2 4 3 5 \u00bc L h \u00f01 Cosh\u00de Cos/ \u00f01 Cosh\u00de Sin/ Sinh 2 4 3 5 (1) According to the fixed position of the driving wires at the top segment of continuum structure, as shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002361_1.4034500-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002361_1.4034500-Figure3-1.png",
        "caption": "Fig. 3 Steam test rig",
        "texts": [
            " For this new system, two main requirements have been formulated. The first one is that the new design should include a concept for interchangeable running surfaces. So, different tribopairs between brush seal and rotating counter surface can be investigated without changing the whole rotor system. Second, thermocouples have to be included in the area of brush seals\u2019 running positions for temperature distribution measurements. A complete schematic overview of the steam test rig, including the new rotor system, is shown in Fig. 3. As mentioned above, the general operation principle has not been changed compared to former setups. The central element is the high-pressure (HP) chamber, consisting of an annular casing enclosed by cover plates at each end. Designed as a double-flow test rig, two opposing brush seal configurations, single or multistage, find space within the high-pressure chamber. Figure 3 illustrates the schematic view of two multistage arrangements. The investigated seal arrangements will be introduced separately. The casing position can be varied relative to the rotor position in order to actively control the seal\u2019s clearance/interference level. The rotor drive concept includes an 80 kW electric motor which is joined to the rotor via a curved teeth coupling. The nominal shaft speed is 10,000 rpm, leading to a maximum surface speed of nearly 160 m/s. The steam is induced in the center of the HP-chamber and is fed back to a power plant steam header after throttling in the individual seal passages"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003409_j.engfailanal.2020.104863-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003409_j.engfailanal.2020.104863-Figure10-1.png",
        "caption": "Fig. 10. Fracture surface of a failed bolt.",
        "texts": [
            " The 11th bolt in 4th row was broken when the load cycles at 143427th cycle of the 1st level of the 6th block (5960 h). It is worth noting that the actual load cycles of bolts were relatively in accordance with the predicted load cycles as shown in Table 5. As observed in the aforementioned analysis, the bolts on both the sides in 4th row failed prematurely. Therefore, the bolts on both sides in 4th row should be designed such that it can withstand fatigue to satisfy the usage requirements. The fracture surface of bolt in 4th row and, 1st column is shown in Fig. 10. The fracture surface of broken bolt was examined by SEM [26]. It was observed that the cracks originated from the joint of the bolt head and bolt shank, and there was a multi-source crack propagating inward. Finally, the bolt broke completely. The morphology of the fatigue crack source region at 200x was observed. As a result of severe stress concentration, the feature of Table 2 Mechanical properties of the steels. Steel grade Elastic modulus (MPa) Yield stress (MPa) Tensile stress (MPa) Q345 \u00d72"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003993_12.2184834-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003993_12.2184834-Figure10-1.png",
        "caption": "Figure 10. Location of instantaneous center of rotation (ICR).",
        "texts": [
            " 9528 95280L-9 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 09/13/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx Then the position of instantaneous center of rotation (x0, y0) is given by:  \u00a0 \u00a0\ud835\udc65! = \ud835\udc3a! \ud835\udc65! + \ud835\udc52!! \u2212  \u00a0\ud835\udc3a! \ud835\udc65! + \ud835\udc52!! \ud835\udc3a!\"  \u00a0 \u00a0 (8) \ud835\udc66! = \ud835\udc3a! \ud835\udc66! + \ud835\udc52!! \u2212  \u00a0\ud835\udc3a! \ud835\udc66! + \ud835\udc52!! \ud835\udc3a!\" Where G1, G2, G12 are given by: \ud835\udc3a! = \ud835\udc65! \ud835\udc65! + \ud835\udc52!! + \ud835\udc66! \ud835\udc66! + \ud835\udc52!! \ud835\udc3a! = \ud835\udc65! \ud835\udc65! + \ud835\udc52!! + \ud835\udc66! \ud835\udc66! + \ud835\udc52!! \ud835\udc3a!\" = \ud835\udc65! + \ud835\udc52!! \ud835\udc66! + \ud835\udc52!! \u2212 \ud835\udc65! \ud835\udc65! + \ud835\udc52!! \ud835\udc66! + \ud835\udc52!! The geometrical location of ICR is illustrated on figure 10. The proposed algorithm is summarized in table 3. Proc. of SPIE Vol. 9528 95280L-10 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 09/13/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx A fast algorithm for estimation of the instantaneous center of rotation for a skid-steered robot equipped with a video camera was developed. The algorithm is based on Horn\u2013Schunck variational optical flow estimation method. To reduce the computational cost of estimation of the smoothness term a filtering in the frequency domain is used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure2-1.png",
        "caption": "Fig. 2 Modified symbol of Wolf with three external shafts and their torques",
        "texts": [
            " x1 rel T1 [ 0\u2014the relative power is transmitted from the sun gear 1 through the planets 2 to the ring gear 3, or in other words, the sun gear appears to be the driving one in the relative motion; \u2022 when x1 rel T1\\0\u2014the relative power is passed from the ring gear 3 to the sun gear 1; In relatively rare cases this situation is not that clear. Then it is appropriate for determining the direction of the relative power to use the method of the samples (Seeliger 1964). Along with the foregoing known principles of mechanics, the following two small but very useful innovations are made: 5. For the greatest possible illustration of the method the known symbol of the Wolf (1958) (Fig. 2) is used, but modified (Arnaudov 1984, 1996; Arnaudov and Karaivanov 2001, 2005a, b, 2010; Karaivanov 2000; Karaivanov and Arnaudow 2002). The each elementary (single-carrier) gear train, which has three external shafts is depicted with a circle, from which the three shafts are coming out, marked differently from the original\u2014i.e. not with numbers, letters or inscriptions, but by the thickness of the lines corresponding to the size of the external torque, as already mentioned above. This little innovation concerning the manner of marking the shafts makes the method very clear and useful"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001836_iecon.2015.7392422-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001836_iecon.2015.7392422-Figure6-1.png",
        "caption": "Fig. 6. Kinematics.",
        "texts": [
            " \u03b8 = atan2 (\u221a 0R2 E,32 + 0R2 E,32, 0RE,33 ) (17) \u03c8 = atan2(0RE,32, 0RE,31) (18) \u03c6 = atan2(0RE,23, 0RE,13) (19) Where, 0RE,ij is the entry in the i-th row and j-th column of the matrix 0RE . The Jacobean matrix Jr\u0307 for convert the joint space angle velocity q\u0307 to workspace velocity r\u0307 can be calculated from (20) Jr\u0307,i = \u2202r \u2202qi (20) It is the general calculation for Jacobean matrix. On the other hand, the Jacobean matrix can also calculated form the dynamics of each joint as Jr\u0307 = [0Rz1\u00d70PE,1 0Rz2\u00d70PE,2 \u00b7 \u00b7 \u00b7 0Rz5\u00d70PE,5] (21) Here, the vector 0PE,i is translation vector shown in Fig. 6(a) and (22). And the vector 0Rzi is third member of homogeneous transformation form joint space \u03a30 to joint space \u03a3i as shown in (28). 002168 0PE,i = 0PE \u2212 0Pi (22) 0Ti = [ 0Rxi 0Ryi 0Rzi 0Pi 0 0 0 1 ] (23) The calculation result of (20) and (21) will be same, and the calculated jacobian matrix is shown as below. 0Rz1\u00d70PE,1 = \u23a1 \u23a2\u23a2\u23a2\u23a3 \u2212l1S1C2+l2S1S2+l3(C1C4+C3S1S4) +l4(C5S1S3\u2212(\u2212C3C4S1+C1S4)S5) l1C1C2+l2C1S3+l3(S1C4+C3C1S4) +l4(\u2212C5C1S3\u2212(\u2212C3C4C1+S1S4)S5) 0 \u23a4 \u23a5\u23a5\u23a5\u23a6 (24) 0Rz2\u00d70PE,2 = [\u2212l1C1S2 \u2212l1S1S2 \u2212l1C2 ] (25) 0Rz3\u00d70PE,3 = \u23a1 \u23a2\u23a2\u23a2\u23a3 \u2212l2C1S3+l3S3C1S4 +l4(\u2212C5C1C3+S3C4C1S5) \u2212l2S1C3+l3S3S1S4 +l4(\u2212C5S1C3+S3C4S1S5) l2S3+l3C3S4+l4(C5S3+C3C4S5) \u23a4 \u23a5\u23a5\u23a5\u23a6 (26) 0Rz4\u00d70PE,4 = [\u2212l3(S1C4+C3C1S4)\u2212l4(S1C4\u2212C3S4C1)S5 l3(C1S4\u2212C3S1C4)\u2212l4(C1C4+C3S4S1)S5) l3S3C4+l4S3S4S5 ] (27) 0Rz5\u00d70PE,5 = [ l4(S5C1S3\u2212(C3C4C1+S1S4)C5 l4C1C2+l2C1S3+l3(S1C4+C3C1S4) l4C1C2+l2C1S3+l3(S1C4+C3C1S4) ] (28) The Jacobean matrix J\u03c9 for convert the joint space angle velocity q\u0307 to workspace angle velocity (\u03b8\u0307, \u03c8\u0307, \u03c6\u0307) can be als calculated from dynamics of each joint as . J\u03c9 = [ 0Rz1 0Rz2 0Rz3 0Rz4 0Rz5 ] (29) The robot arm used in this paper can control two angle (\u03c8, \u03c6) shown in Fig. 6(b). Therefore the Jacobean matrix J \u2032 \u03c9 for convert the joint space angle velocity q\u0307 to workspace angle velocity (\u03c8\u0307, \u03c6\u0307) can be calculated from as J \u2032 \u03c9 = [ 0 C1 C1 \u2212S1S2+3 C1C4+S1S2+3S4 1 0 0 \u2212C2+3 \u2212S2+3S4 ] (30) The last degree of freedom \u03b8 is the rotation of drill. In this subsection the simulation and experiment of position tracking was conducted. First, the inverse kinematics was calculated for the position calculation. Next, tow position tracking simulation was explained considering the minimum norm and angle posture"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure6.17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure6.17-1.png",
        "caption": "Fig. 6.17 Measurement platform of the force output",
        "texts": [
            " . . . . . 109 Figure 6.12 Magnetic field variation versus h at r \u00bc 12:5 mm, z \u00bc 20 mm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110 Figure 6.13 Tubular linear machines with Halbach array . . . . . . . . . . . . 111 Figure 6.14 Magnetic field distribution at z \u00bc 0 mm . . . . . . . . . . . . . . . 112 Figure 6.15 Magnetic field distribution at r \u00bc 12 mm . . . . . . . . . . . . . . 113 Figure 6.16 Force comparison of different magnet arrays . . . . . . . . . . . . 114 Figure 6.17 Measurement platform of the force output. . . . . . . . . . . . . . 114 Figure 6.18 Force output of experimental versus analytical results. . . . . . 115 Figure 6.19 Experimental study on armature reaction field . . . . . . . . . . . 116 Figure 6.20 Armature reaction field versus z0 for a three-coil winding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117 Figure 6.21 Armature reaction field versus z0 for a five-coil winding . . . . 117 Figure 6.22 Schematic diagram of inductance measurement ",
            " It is found that force output of the linear machine with dual Halbach array approaches to twice of that of interior Halbach array and more than 30 % higher than that of exterior Halbach array. Both magnetic field and force comparisons show that linear machines with dual Halbach array have a better performance. Therefore, the dual Halbach array is promising to be applied in high power density systems like direct drive valve. (a) Br variation (b) Bz variation The measurement of force output is an essential part to analyze system performance [16\u201318]. In order to measure the force output, one measurement platform is developed as shown in Fig. 6.17. The strain type pressure sensor is mounted on output shaft of the mover. It is fixed on a high-precision translational motion stage. The mover can thus be adjusted to measure the force output for different mover positions. (a) Br variation (b) Bz variation The force output is measured within stroke and compared with the theoretical torque model with step size of 1 mm. The measurement result is shown in Fig. 6.18. It is found that the experimental result fits with the analytical model well, which indicates that the precision of the analytical force model is acceptable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002643_robio.2016.7866657-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002643_robio.2016.7866657-Figure9-1.png",
        "caption": "Fig. 9. Clamping simulation: simple spherical soft object.",
        "texts": [],
        "surrounding_texts": [
            "In the second check all nodes that pass the first check are studied: For each evaluated node a ray is launched along the normal opposite direction; if this intersects a triangle of the soft object and the scalar product between the normal of the two triangles involved in the collision is less then zero, then the current node is correctly colliding.\nIn detail, in Fig. 4 three steps of the algorithm are shown. In the first step, the two objects are not colliding. In the second step, only the bottom nodes of the rigid object are involved in the collision. The most critical case is evaluated in the third step in which the rigid object is immersed in the soft object and both the bottom and the top nodes are involved in the collision. In this latter case, the top nodes must be excluded from the collision computation.\nFinally, in order to update the soft object shapes, the elastic force\nf = kt (2)\nis applied to all the nodes of the soft object that are involved in the collision. In Eq. (2), t is the distance computed with the ray casting procedure. This force causes a deformation of the soft object in order to bring it outside the volume of the rigid body.\nThe rigid body used in [1] is a parallelepiped and the nodes used for collision detection are placed only on the two larger surfaces. The symmetrical shape of the body allows an easy manual placement of the nodes by defining the resolution along the planar directions of the two surfaces, see Fig. 5 (left).\nIn this work, the proposed algorithm has been extended from rigid boxes to generic convex objects modelled with triangular meshes. The algorithm in [1] requires that nodes on parallel surfaces are associated in pairs that belong to the same surface normal. For convex objects, as in Fig. 5 (middle), the nodes association is realized by placing a node in the centroid of each triangle mesh and by associating it to a node that belongs to the triangle intersected by the ray casting. The ray source is the node itself and the direction is the opposite of the normal to the triangle to which the nodes belongs. In this way, each triangle finds its opposite pair.\nWith respect to [1], where if a node x is associated to a node y also the vice-versa holds, in this case the nodes are not always mutually associated. This is not critical for the application of the algorithm , but some problems may arise.\nThe rigid object represented in Fig. 6 (right) presents triangular faces in close proximity to the side edge, thus, very close to one another. Indeed, the distance between two triangles on the side edge may be of two orders of magnitude smaller than the thickness of the object, as in Fig. 6 (right).\nMoreover, as an effect of triangle orientation, if a node on the top left, that belongs to a triangle in proximity to the edge side, casts a ray in the opposite direction of its normal, a node on the bottom right can be associated, as represented in Fig. 6 (left).\nThis can be critical for the proper application of the algorithm. Indeed, when the symmetrical arrangement of the associated nodes is lost, the wrong triangle associations can cause unfeasible deformations. To solve the problems arising for these critical nodes, the ray casted has no more infinite extension but is limited according to the dimensions of the scene and especially of the colliding soft bodies. In this way, infeasible collisions are highly reduced. Moreover, since there is no mutual association between the nodes, it is not possible to study the state of a node by verifying only the state of the opposite one. When a node belongs to an infeasible pair, the associated index is stored in a vector collecting all the indices of the nodes that cannot collide in the current simulation step. Thus, a cross-check is necessary at each step.\nIn robotic surgical simulation, it is very important to have a precise control of the tools involved in the simulation but could be equally important to have the possibility to sense the scene through haptic force feedback. In fact, although in most",
            "surgical robots the force feedback was not yet implemented, this technology could be helpful in the process of training in order to create force guided simulations, to provide a more realistic perception of the simulated scene and to help the surgeon to create a mental link between the applied forces and deformations perceived through the visual feedback.\nIn our work a Novint Falcon 3D haptic device is used in order to improve the sense of realism during the simulation. The contact force has been modeled using the Hunt Crossley model (Eq 3) presented in [3] incorporating a non linear spring in parallel with a non-linear damper to model the viscoelastic dynamics:\nF (t) =\n{ kxn(t) + \u03bbxn(t)x\u0307(t) x(t) \u2265 0\n0 x(t) < 0 (3)\nwhere x is the deformation, k and \u03bb are the elastic and viscous parameters of the model respectively, and n is a real number (usually close to one), that takes into account the geometry of the contact surfaces.\nAt each time step, the deformation x is the distance between the current centroid of a triangle and the centroid of the same triangle at time zero. For this purpose, during the object initialization, the original position of each triangle vertices is stored. Obviously, the calculation of x is executed if and only if at least one of the nodes belonging to the triangle is in a colliding state. The total force is computed by adding the components related to each deformed triangle involved in the collision.\nIn the first case study, a single rigid object (like a spatula) interacting with a deformable sphere has been considered. The sphere is composed by 128 triangular meshes, and is placed on a ground plane. The position of the spatula in the 3D space is given by the Novint Falcon 3D haptic device. The visual rendering has been realized using OpenGL.\nIn the second case study, the scene is constituted of a clamp realized with two collidable rigid objects and a handle not collidable but with solely aesthetic functions. The haptic interface allows three degrees of freedom (DoFs) for the controlled object motion. However, the clamp requires one DoF for closing and opening. Therefore, for simulation purposes, the motion of the clamp has been limited to a plane and the extra DoF has been used to control the closing/opening on different object shapes, such as a sphere and an elongated\nobject shape, that simulate more realistically vessels and, in general, organic tissues.\nBy applying the default collision detection method proposed by Bullet Physics, it follows that the spatula collides with the soft body in a consistent way only for small displacements and slow motion. Fast displacement of the spatula can cause a complete penetration into the soft body surface without further collisions, see Fig. 7 (left).\nMoreover, the application of only the ray casting method without node association (second step of the algorithm) causes drawbacks if the dimension of the rigid body along the collision direction is not large enough. The collision appears unstable, as in Fig. 7 (right), when both the nodes belonging to opposite surfaces apply collision forces to the soft body even if they are unfeasible.\nThis drawback is overcome by applying also the second step of the algorithm, that detects the unfeasible collisions and excludes them from the study at the current simulation step.\nThe algorithm has been tested with convex rigid objects and with a clamp grabbing deformable objects, see Figs. 8,9,10.",
            "The results are realistic both in terms of visual rendering and force feedback, see Fig. 12.\nIn Fig. 11 the computation time of the algorithm as a function of the soft mesh triangles number is shown. A good performance is obtained also with a complex soft object (more than 4000 triangles) and using a basic hardware platform (Intel I5 processor, Nvidia GTX750M VGA, 4GB of RAM).\nIn the light of the promising results with respect to the Bullet default collision detection algorithm, future work\nThis research has been partially funded by the EC Seventh Framework Programme (FP7) within RoDyMan project 320992. The authors are solely responsible for its content. It does not represent the opinion of the European Community and the Community is not responsible for any use that might be made of the information contained therein.\n[1] A. Fukuhara, T. Tsujita, K. Sase, A. Konno, X. Jiang, S. Abiko, M. Uchiyama, \u201cProposition and evaluation of a collision detection method for real time surgery simulation of opening a brain fissure\u201d, ROBOMECH Journal, vol. 1, no. 6, 2014. [2] URL: https://bulletphysics.org/. Bullet phisics Web page. [3] F. Ficuciello, R. Carlon, L.C. Visser, S. Stramigioli, \u201dPort-hamiltonian modeling for soft-finger manipulation\u201d, IEEE/RSJ International Conference on Intelligent Robots and Systems, pp. 4281\u20134286, 2010. [4] C. Ericson, Real-time collision detection, CRC Press, 2004. [5] S. Kockara, T. Halic, K. Iqbal, C. Bayrak, R. Rowe, \u201cCollision\ndetection: A survey\u201d, IEEE International Conference on Systems, Man and Cybernetics, 2007. [6] T.J. Purcell, I. Buck, W.R. Mark, P. Hanrahan, \u201cRay tracing on programmable graphics hardware\u201d, ACM Transactions on Graphics, vol. 21, pp. 703\u2013712, 2002. [7] T. Moller and B. Trunbore, \u201cFast, minimum storage ray-triangle intersection test\u201d, Journal of graphic tools, vol. 2, pp. 21\u201328, 1997. [8] URL: https://www.sofa-framework.org/. SOFA Framework Web page."
        ]
    },
    {
        "image_filename": "designv11_30_0003511_j.triboint.2020.106696-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003511_j.triboint.2020.106696-Figure9-1.png",
        "caption": "Fig. 9. Schematic diagram of distribution of thermocouples.",
        "texts": [
            " A copper pipe was used to extract the HP-chamber gas and a self-made pressure pipe was used to extract the downstream gas. A pressure scanner was used to measure the pressure, the range of which is 0\u20130.6 MPa. The accuracy of the scanner is better than \u00b10.05% FS. Temperature around the bristle tip was measured by thermocouples. Four thermocouples were evenly distributed along the circumferential direction of the brush seal, and pressed the tip of the front bristles. The diagram of the distribution of thermocouples is shown in Fig. 9, and the J. Fan et al. Tribology International 154 (2021) 106696 critical object is shown in Fig. 10. The measuring range of the thermocouples is from 0 to 1000 \u25e6C. The accuracy of the thermocouples is \u00b10.75% of the temperature to be tested. Photoelectric sensor was used to record the rotating speed \u03a9. The measuring range is 2.5\u201399999 rpm. The distinguishability is 1 rpm (1000\u201399999 rpm). The sampling frequency is 60Hz. Wear will increase the inner diameter of brush seal, Di, and reduce the outer diameter of rotor, Do"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002575_itsc.2016.7795745-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002575_itsc.2016.7795745-Figure1-1.png",
        "caption": "Fig. 1. Tilt-rotor UAV frames.",
        "texts": [
            " Raffo is with the Department of Electronics Engineering, Federal University of Minas Gerais (UFMG), Belo Horizonte, MG, Brazil. 3S. Esteban is with the Department of Aerospace Engineering, Universidad de Sevilla, 41092, Seville, Spain. {sesteban}@us.es timal controller [8], and feedback linearization [9]. However, all these works neglected aerodynamic forces. In this work, in order to improve the forward flight performance, tail controlled surfaces are added to the tiltrotor UAV, as can be seen in Fig. 1. Although the flight performance improvement is achieved with vertical tail plane (VTP) and horizontal tail plane (HTP), the control design becomes challenging. This is because when the tilt-rotor UAV is in helicopter-flight mode (low velocity) the deflection of rudder or elevator do not produce significant effects, as opposed to in forward flight, where small deflections produce significant aerodynamic forces that result in an increased control effectiveness in both longitudinal and lateral motion",
            " This adaptive controller is extended here using the system statespace representation and assuming the stability guarantee of the mixed H2/H\u221e robust controllers into polytopes. This section develops the tilt-rotor UAV dynamic model using the Euler-Lagrange formulation. The tilt-rotor UAV considered in this work is formed by three bodies. The main body is composed by the fuselage, horizontal and vertical stabilizers. The reminder two bodies are each of the two thrusters\u2019 groups (servomotors with rotors) interconnected to the main body by revolute joints. In order to obtain the forward kinematics, seven frames are defined, as shown in figure 1: the inertial frame I; the body frame B; the frames rigidly attached to the main body center of mass, C1, and to the centers of mass of the two thrusters\u2019 groups, C2 and C3, correspondent to right and left sides, respectively, apart from the auxiliary frames Caux 2 and Caux 3 . The homogeneous transformation matrix from B to I is expressed as HIB = [ RIB dIB 01\u00d73 1 ] , where dIB , [x y z] \u2032 is the position of the origin of frame B with respect to I, and RIB , Rz,\u03c8Ry,\u03b8Rx,\u03c6 is the rotation matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001898_s00542-015-2527-2-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001898_s00542-015-2527-2-Figure3-1.png",
        "caption": "Fig. 3 Air bearing stiffness model, showing the linear springs used to model the force between the slider and the disk",
        "texts": [
            " The slider and disk are connected by air bearing springs. A real air bearing system can be described as a nonlinear spring in the operational shock situation. However, the air bearing spring can be simplified as a linear spring to speed up the simulation of the structural dynamic response (Liu et al. 2009). Considering Liu et al. (2009), we connected the slider and disk surface with four linear springs located at the centers of the trailing (TE) and leading (LE), inner (IE), and outer (OE) edges, as shown in Fig. 3. The HDD FEM model is constructed as a \u201ctied\u201d type model, designed so that the upper and lower ends of the shaft are fixed to the base and cover with screws. When the shaft is fixed, the structure has high structural stiffness and high resonance frequency. Therefore, the cover is connected to the upper end of the spindle, pivot bearing, and six bolting positions in the base in the FEM model. During a shock, ramp\u2013disk contact occurs. A nonlinear contact model was constructed to consider ramp\u2013disk contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.18-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.18-1.png",
        "caption": "Figure 2.18 Suction electrodes: (A and B) components of a suction electrode and (C) variation of suction electrode with a recessed Ag AgCl disc and rubber dome.",
        "texts": [
            " In addition to gold, silver, and stainless steel, they can be made of a metallic alloy of copper, zinc, and nickel, known as German silver or alpaca. German silver is extensively used because of its hardness, toughness, resistance to oxidation, and high electrical resistance; the percentage of the three elements varies, ranging for copper from 50% to 61.6%; for zinc from 17.2% to 19%; and for nickel from 21.1% to 30%. The suction electrode has a metallic dome shape with a connector to the electric cable, and a suction rubber bulb (Figure 2.18A and B). Different metal and alloys can be used in the dome: gold, silver, Ag AgCl, Ni Cu, and german silver (an alloy of copper, nickel, and zinc; for instance, 60% copper, 20% nickel, and 20% zinc). The main use of suction electrodes is in precordial ECG recording. The rubber bulb is squeezed and then the electrode is connected to the chest through suction, which removes the air between the electrode and the skin, keeping the electrode in place without the need of elastic bands or adhesives. An electrolyte layer is applied in the edge of the dome before the electrode is adjusted to the skin surface. The suction can cause skin irritation and this type of electrode should be used for short periods of time; moreover, it also causes discomfort to the patient since its use can be painful. It is manufactured in different diameters to child and adult use (15, 24, and 30 mm). Among the metal electrodes, the suction electrode is the one with the higher impedance due to the small contact area (edge of the dome). Figure 2.18C shows a suction electrode that has the dome made of rubber with a recessed metal disc inside (Ag AgCl), which contributes to decrease the electrode impedance (larger contact surface) and to reduce the discomfort (rubber, instead of metal suctions the patient skin). The electrolytic gel layer is applied on the disc surface. One example of this version of suction electrode, also used in chest wall for ECG recording, is the H5 1800 fabricated by Servopraxs. This type of electrode is formed by a metal disc with a wire or a cable connector welded to the center of one of the disc faces (Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002092_j.mechmachtheory.2016.04.003-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002092_j.mechmachtheory.2016.04.003-Figure2-1.png",
        "caption": "Fig. 2. Radius vectors of a mass particle dm on a rigid body.",
        "texts": [
            " In paper [1,2], considering the rotational equilibrium of the lower part and upper part of the leg in the fixed leg frame, Euler's equation gives \u2212mdrd ad \u00femdrd g\u2212IdA\u2212W IdW \u00feMus\u2212r Fp\u2212Mp\u2212CuW \u00bc 0 \u00f01\u00de \u2212muru au \u00femuru g\u2212IuA\u2212W IuW \u00fe S Fs \u00fe r Fp \u00feMp\u2212Cs W\u2212\u03c9\u00f0 \u00de \u00bc 0 \u00f02\u00de where Fp is the vector force at the prismatic joint exerted by the lower part on the upper part acting at a point r, Mp is the vector moment at the prismatic joint acting on the upper part, other parameters refer to Fig. 1. The basic error of these equations is that the angular momentum theorem isn't properly applied. Let O0 be a point fixed in inertial space, C is the center of mass and B is an arbitrary point fixed on the body (see Fig. 2). In combination with the vector operation rules and Newton's axioms, the angular momentum theorem for a rigid body is obtained in the final form m\u03c1c \u20acrB \u00fe JB _\u03c9\u00fe\u03c9 JB \u03c9 \u00bc MB \u00f03\u00de 3 in which \u03c9 is the absolute angular velocity of the body, \u20acrB, JB and MB are the absolute acceleration, inertia tensor and resultant moment about point B respectively. In particular, if reference point B is chosen as the body center of mass (\u03c1c=0) or body-fixed point B which is also fixed in inertial space (\u20acrB \u00bc 0), the Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000834_2014-01-2064-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000834_2014-01-2064-Figure3-1.png",
        "caption": "Figure 3. The simplified 8 DOFs model",
        "texts": [
            " The drive shaft and half shaft are discretized by finite element method to generate the modal neutral file (MNF) which is imported into the dynamic model that is connected with adjacent rigid components by kinematic pairs. The rigid and flexible coupling model is established by combining the dynamic model and the rigid components, as shown in Figure 2. In order to reduce the complexity of the DOE analysis and parameter sensitivity analysis, the multi-body dynamic model is simplified to an 8-DOF model, as shown in Figure 3. The symbol's mean and parameters' value of system are shown in Table 1 and Table 2 respectively. The inertia of transmission shafts can be equated as follow, i.e. J\u2032 is equivalent value of J, as shown in figure 4, (1) Assume \u03c91/\u03c92 = i (2) The relation between the inertia and equivalent inertia is: (3) The torsional stiffness for drive shaft and half shaft can be calculated. For solid shaft, the stiffness is: (4) For hollow shaft, the stiffness is: (5) Where G is shear modulus; D and d are shaft outer and inner diameters respectively; and l is shaft length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002404_mmar.2016.7575286-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002404_mmar.2016.7575286-Figure2-1.png",
        "caption": "Fig. 2. Real images Pc1, Pc2 and ideal images Pci1, Pci2 of the point P",
        "texts": [
            " The reference frame xyz is associated with the robot station. The F1 is focal of the Camera1, and F2 - focal of the Camera2. The Pcor is the point calculated from the coordinates of the images Pc1 and Pc2 of the characteristic point P. These images are read properly from the Camera1 and Camera 2. The real images Pc1 and Pc2 are different from the ideal images Pci1 and Pci2. Points Pci1 and Pci2 lie on a straight passing through the point P and respectively by points F1 or F2. These points are illustrated in Fig.2. 1This work was supported by Silesian University of Technology Grant BK-227/Rau-1/2015/2. 978-1-5090-1866-6/16/$31.00 \u00a92016 IEEE 1069 The differences between the real and ideal images are caused by optical errors of cameras. With cameras streaming video the server downloads the images and writes them to disk. Then the server processes the markers images to black and white form. The background color is black and the white marker. Figure 3 illustrates an example of an processed image. From so processed image the coordinates of markers geometric centers are calculated in the frame associated with the matrix sensor corner of the camera, in pixels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003632_j.ymssp.2020.107312-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003632_j.ymssp.2020.107312-Figure1-1.png",
        "caption": "Fig. 1. The planetary rover.",
        "texts": [
            " The effectiveness of the proposed method and models is demonstrated by experiments with a physical rover protype. The rest of this paper is organized as follows. Section 2 describes the model of the six-wheeled planetary rover. The coordinated tracking controller is designed in Section 3. In Section 4, the experimental results are represented. This paper ends with the conclusion in Section 5. The six-wheeled planetary rover under investigation is customized to imitate the Yutu lunar rover. The rover consists of a chassis, two pairs of suspensions and six rigid wheels with lugs on them, as shown in Fig. 1. The Sojourner structure is adopted by the rover. This mechanical structure refers to a rocker-bogie suspension structure, which connects the wheels and the chassis. There are no elastic elements and actively driven motors in suspension. The rover usually contains two symmetrical suspensions. Six wheels are actively driven and four steering motors are arranged on four corner wheels. Define (x, y, z)T denotes the positions of the center of the planetary rover in the world coordinate frame. When moving on the rough terrain shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000918_icems.2013.6754383-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000918_icems.2013.6754383-Figure9-1.png",
        "caption": "Fig. 9. Stress distribution at 10,000 rp",
        "texts": [
            " The motors are controlled by the field weakening method. The maximum output of the motor proposed shows 15.1 kW and that of Type A is 15.4 kW as shown in Fig. 7. The output of the motor proposed decreases due to the torque reduction. Fig. 8 shows the relation between the phase angle \u03b2 and the rotor speed. As the rotor speed becomes higher, the phase angle \u03b2 becomes larger to maintain a line voltage of 400 Vrms as shown in this figure. There is almost no difference of \u03b2 between the two motors. C. Mechanical Strength of Rotor Fig. 9 shows the stress distribution at 10,0 in the third layer are made along the normal d motor\u2019s axis as shown in Fig. 2(a). The improves the mechanical strength because th directly press the ribs which are made a direction. The stress at the edge of the magn shown in Fig. 9. The yield stress of iron core is 275 MPa, a maximum stress is calculated by (1). a y sF \u03c3 \u03c3 = where sF safety factor y\u03c3 yield stress a\u03c3 allowable maximum stress When a safety factor is 1.5, the allowable is 183.3 MPa. The maximum speed is d maximum stress in the rotor reaches 183.3 M results of analysis, the maximum speed of the is 10,900 rpm and that of Type A is 8,300 rpm factor is considered. D. Loss and Efficiency for Speed Fig. 10 and 11 show the iron and copper l in the motor proposed and Type A, respect are driven by the field weakening method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003914_gt2015-43940-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003914_gt2015-43940-Figure3-1.png",
        "caption": "Figure 3. CT data and fitted surface of the lower surface of one channel of the M-2x-In coupon (Note: coordinate axes have different scales).",
        "texts": [
            " An in-house code was developed to take the point cloud data and quantify the dimensions and roughness of each channel surface. Each surface of every channel was approximated using a polynomial surface fit. The equations of the fit surfaces were 3rd order in the height and width dimensions and 5th order in the streamwise length dimension. The fitting algorithm utilized a bisquare weighting approach to limit the influence of the outliers on the mean surface location. An example of the surface fit to the data is shown in Figure 3 with an exaggerated scale in the height coordinate to make the variation more visible. To better understand how the surface fits the data, Figure 4 was generated by cutting the surface fit and CT data using the cut plane illustrated in Figure 3. The resulting line plots show the deviation of the roughness. As shown in Figure 4, the polynomial surface fit expresses the mean surface location. Also indicated in Figure 4 was the typical warping of the walls of the coupons, which was \u00b1 80 microns. The fitted surfaces were used to calculate the average manufactured height and width dimensions of each channel. The final height dimension was calculated as the average distance between the top and bottom surfaces, and the final width dimension was calculated in a similar manner using the side surfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure19-1.png",
        "caption": "Fig. 19. Bifurcation of parallelograms and antiparallelograms.",
        "texts": [
            " Since the mechanism is an open-chain mechanism with 3 links and 2 R joints, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration III ( Fig. 17 ), when the axes of R joints J and J coincide, the axes of R joints J and J coincide, and the axes of R joints J and R2 R4 R6 R8 R1 J R5 coincide during the movement, the mechanism makes a planar motion, which is called plane-motion Mode V. Its typical configurations are shown in Fig. 18 . Due to the characteristics of antiparallelogram unit ( Fig. 19 ), it is a bifurcation point at \u03b8AA = 0 or \u03b8AA = \u03c0 . At a bifurcation point, an antiparallelogram could turn into a parallelogram. Therefore, there are some more special configurations as shown in Fig. 20 (no configurations interfere with each other by the layering). In addition, when interference is not considered, some special configurations are shown in Fig. 21 . Some configurations of plane-symmetric Mode I, plane-symmetric Mode II, plane-motion Mode III, plane-motion Mode IV, and plane-motion Mode V can be transited to each other, and they have a certain relationship as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003781_ssrr50563.2020.9292590-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003781_ssrr50563.2020.9292590-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the differential drive mobile robot [3]",
        "texts": [
            " The power is eventually found as the multiplication of the motor voltage and current in the combined model. This model was already developed and extensively tested in a prior work [3], and it is summarized here for easy reference. 1) Dynamic model of a two wheel differential drive robot: The dynamic model for a differential drive mobile robot can be represented by the following differential equation system based on the Lagrange formulation [14], [15], [16], and the kinematic model of the robot, whose schematic is shown in Fig. 1 : \u03bb1 sin\u03c6+ \u03bb2 cos\u03c6 = mx\u0308\u2212mcd(\u03c6\u0308 sin\u03c6+ \u03c6\u03072 cos\u03c6) \u2212\u03bb1 cos\u03c6+ \u03bb2 sin\u03c6 = my\u0308 +mcd(\u03c6\u0308 cos\u03c6\u2212 \u03c6\u03072 sin\u03c6) \u03c41 \u2212 cl\u03bb2 = mccd(y\u0308 cos\u03c6\u2212 x\u0308 sin\u03c6)+ (Ic2 + Iw)\u03b8\u03081 \u2212 Ic2\u03b8\u03082 \u03c42 \u2212 cl\u03bb2 =\u2212mccd(y\u0308 cos\u03c6\u2212 x\u0308 sin\u03c6)\u2212 Ic2\u03b8\u03081 + (Ic2 + Iw)\u03b8\u03082 (1) 2) Dynamic model of a motor: As shown in Fig. 1, a WMR has usually a symmetric structure, therefore both motors that drive the robot can be assumed to be identical, and hence they can be represented by the following differential equation system with the same parameters [17]: Li\u0307+Ri+Kvn\u03b8\u0307 =V Ktni\u2212 Is\u03b8\u0308 \u2212 \u03bd\u03b8\u0307 =\u03c4 + Tm (2) The first expression in 2 represents the voltage equation of a DC motor, and the second expression is the corresponding current equation containing the load torque applied to the 137 Authorized licensed use limited to: Raytheon Technologies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000301_gcc45510.2019.1570521870-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000301_gcc45510.2019.1570521870-Figure3-1.png",
        "caption": "Fig. 3. Presentation of the final prototype design.",
        "texts": [
            " Click on the \u201cstreaming\u201d button (it will run the thermal camera and show the live thermal streaming video). 4. Click on \u201cstop streaming\u201d button to take a capture. 5. Run the \u201cthermal image\u2019s\u201d, call up the thermal capture (thermal photo). 6. Select the interested area (overheating location points) in the thermal photo. 7. Input required parameters related to the throughput capacity calculations (for more detail check Section II). 8. Calculate the throughput current of the power line. Let us explain the operating procedure of the final prototype of a proposed system. The final prototype presented in Fig. 3. In order to perform the operation process of the proposed simplified inspection system, the specific drone used (see the specification above). The function of the drone is to carry the designed electrical circuit to the specific interested area (examined OHL location) nearby a phase conductor in order to ensure data capturing (safety distance ensured). Raspberry Pi module B connects to the thermal camera and transmits the thermal photo from this camera to the ground server (user). It has a Wi-Fi range of 50 m, which is an optimum solution since phase conductors of the TL located at quite high elevations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001551_elk-1207-55-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001551_elk-1207-55-Figure9-1.png",
        "caption": "Figure 9. Electrical circuit of the DC motor.",
        "texts": [
            " It is obvious that vector \u03b8 tends to fixed values. The final value of \u03b8 is [\u22121.6511\u22121.11633.4305], which is considered to be an estimation of our ideal \u03b8\u2217. It is easy to observe that matrices P and \u03b8\u2217 satisfy (7). The control input is smooth and limited and is presented in Figure 8. The results show that the tracking error converges to 0 while adjusting gains adaptively. 0 1 2 3 4 5 6 7 8 9 10 -2 -1 0 1 2 3 4 5 Convergence of adjustable parameters (vector \u03b8) Time (s) \u03b8 \u03b81 \u03b82 \u03b83 Figure 7. Convergence of the adjustable parameters (vector \u03b8) . Figure 9 shows a simple DC motor. Since typically La is small, if we neglect the armature inductance, we can write the system equations as [17]: { Va (t) = Ria (t) +Kw\u03c9 (t) J d\u03c9 dt = Ktia (t)\u2212 b\u03c9 (t) , (31) where ia (t) , Va (t) , R are, respectively, the armature current, voltage, and resistance. J is the moment of inertia. Kt,Kw are, respectively, the torque and the back electromotive force, developed with a constant excitation flux. The constant b is the damping coefficient and \u03c9 (t) is the angular speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001076_s12555-013-0057-1-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001076_s12555-013-0057-1-Figure13-1.png",
        "caption": "Fig. 13. Collaborative handling of an object by three wheeled mobile manipulators.",
        "texts": [
            " Although rather sharp rises of error are observed when the steering angle changes much, the tracking performance is in general satisfactory. Concerning the manipulability, it approaches its maximum level (for \u03b3 = 0), as time goes on. 5.2. Collaborative task by differential-drive wheeled mo- bile manipulators In this simulation, we show how to conveniently transform three collaborative WMMs into an equivalent parallel manipulator system with articulated legs through the virtual joint based modeling. We consider a simple example where a common plate is carried by the three mobile manipulators as shown in Fig. 13(a). All the mobile platforms are assumed to be differential-drive type, and the robot arms, two one-link arms and one twolink arm, are attached on top of the mobile platforms one by one. If these WMMs are transformed into equivalent articulated manipulators via the virtual joint modeling, the overall system looks like a common parallel manipulator having three legs of two R-P-R-R and one R-P-R-R-R types. Now, we can simply rely on previously studies on parallel manipulators to understand Hyunhwan Jeong, Hyungsik Kim, Joono Cheong, and Wheekuk Kim 1068 the current collaborative system",
            " Among the joints, only those conjoining the robot arms and the plate are assumed to be passive. Our strategy for this task is, then, to apply the inverse kinematic scheme to each WMM separately using (16) since each joint location in the plate can be considered as the end-effector of the corresponding manipulator arm. Here, we do not consider the secondary objectives of the inverse kinematics because our purpose of simulation is just to demonstrate the efficacy of the virtual joint based modeling. Fig. 13(b) shows the traces of the WMMs and the plates with the final and some intermediate configurations. (Note that the initial posture is given in Fig. 13(a)). The tracking error was negligible and the desired circle trajectory was well followed. Although the simulation task was plain, we could see how the system of collaborative WMMs is effectively conceptualized as a parallel manipulator system, which is familiar, via the virtual joint based kinematic modeling. Thus, further sophisticated research issues of the original system, related to kinematics, statics and control, could be investigated by borrowing theoretical results from parallel manipulators"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003035_012007-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003035_012007-Figure4-1.png",
        "caption": "Figure 4. The PRSM failure test bench.",
        "texts": [
            " Grease failure will aggravate the wear, increase the heat at work, cause a decline in PRSM precision, produce deformation, and eventually lead to the entire PRSM break. In a word, grease failure usually appears before the PRSM whole fracture. Therefore, reasonable judgment on whether grease in the working process is present in PRSM effectively guarantees the PRSM running safety. And it can timely remind the operator whether to add grease to PRSM and provide certain judgment basis. Figure 3 indicates the working state of PRSM with or without grease. Figure 4 shows the PRSM failure test bench, and it consists of four parts, motor, reducer, PRSM and hydraulic load system. The motor drives the reducer, which drives the PRSM to run. And the PRSM pushes the hydraulic load to reciprocate. The vibration acceleration sensor is located at the upside of the nut and the sampling frequency is 20480Hz. Table1. The working condition of PRSM. Working state speed\uff08r/min\uff09 load\uff08KN\uff09 40 3.5 With grease 6 9 3.5 Without grease 6 9 Three working conditions are shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002508_s11432-015-0153-2-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002508_s11432-015-0153-2-Figure4-1.png",
        "caption": "Figure 4 Flight experiment stand.",
        "texts": [
            " Here, we choose \u03c91 as the virtual control input of (34), and define the pseudo control v\u0398 as v\u0398 = \u0398\u0307r + k\u0398(\u0398r \u2212\u0398), (35) where \u0398r is the reference signal of \u0398, k\u0398 is a positive matrix. Then, the desired angular rate \u03c91r can be selected as \u03c91r = A\u22121 \u0398 (v\u0398 \u2212B\u0398r). (36) Two flight experiments have been carried out to investigate the control performance of the proposed scheme on the real helicopter. To facilitate experiment implementation and to guarantee the safety of the helicopter, we construct a five DOFs flight experiment stand in this paper [33], as shown in Figure 4. The stand is a mechanical construction able to hold a helicopter, allowing basic movements while protecting it from damaging and crashing. The roll, pitch and yaw axes provide the whole DOFs of rotation, and the main and elevation axes provide two DOFs of translation. The helicopter is fixed on the stand when flying and the influences of the stand are treated as disturbances which can be estimated by the ESO in this paper. For safety, all the external commands in the experiments are generated by the pilot through RC controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002806_s12206-020-0115-6-Figure12-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002806_s12206-020-0115-6-Figure12-1.png",
        "caption": "Fig. 12. Physical condition at the time of deposition using (a) orientation FWFNA; (b) orientation BWF.",
        "texts": [
            " 9 and 11, the residual stress value was more compressive at higher values of hardness in the deposited material. According to literature, compressive residual stress decreased gradually with increasing hardness of the material [18]. Moreover, the hardness value increased simultaneously with decreasing residual stress. Thus, more compressive residual stress and higher hardness value in the component is beneficial to obtain a higher fatigue life. Also, the residual stress in FWFNA orientation was more compressive than other samples. This observation can be explained using Fig. 12. When the metal deposition was conducted using FWFNA orientation, the deposition direction and the direction of the feeding wire were opposite each other, as shown in Fig. 12(a). The region \u201cA\u201d represents deposited material, \u201cB\u201d represented the molten material, and \u201cC\u201d represented the material coming out of the wire feeding nozzle to be melted. Hence, due to the motion of the feeding wire in the opposite direction, the material in region C exerts force on the region B (i.e., molten material). Considering this compressive force, more compressive residual stress was induced in FWFNA sample than any other orientation sample. Alternatively, in BWF orientation (Fig. 12(b)), wire feeding direction and deposition direction are the same. Thus, region B is not subjected to as much compressive force as in FWFNA. The microstructure analysis for all four samples was done at the points indicated in Fig. 8. Five points equidistant to each other were selected for this analysis. The microstructure analy- sis was conducted to obtain the volume fraction in the deposited material (ASTM-E-124503). The grain size measurements (ASTM-E-112-13) were obtained for all four samples"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003181_tte.2020.3004734-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003181_tte.2020.3004734-Figure8-1.png",
        "caption": "Fig. 8. Torques caused by rotor position estimation errors during the d-axis current pulse injection. (a) Positive estimation error. (b) Negative estimation error.",
        "texts": [
            " Downloaded on July 15,2020 at 14:27:32 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. According to the characteristics of the investigated VFMM, the MS of the machine can be elevated by injecting a positive d-axis current pulse, which has the capability to lock the rotor position. As shown in Fig. 8, if there is an estimation error of the estimated d-axis, the injected positive d-axis current will generate a torque to make the actual d-axis align to the estimated d-axis, resulting in a decrease of estimation error and finally keeps the rotor still. Thus, the rotor position can be considered as constant during the short period of the d-axis current pulse injection. The measured relationship between the PM flux linkage and the magnitude of d-axis current pulse is stored in a lookup table as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003185_012049-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003185_012049-Figure9-1.png",
        "caption": "Figure 9. Flux distribution in faulty motor (43%). (a) Flux lines. (b) Flux density.",
        "texts": [
            "1088/1757-899X/872/1/012049 The flux distribution of a fine fettle IM appears in Figure 8. It is seen from Figure 8(a), the dispersion of magnetic lines over the solid motor is balanced over the post pitch and each shaft is found at an ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012049 IOP Publishing doi:10.1088/1757-899X/872/1/012049 attractive pivot of 360\u02da/p geometrical degrees, where p is the number of posts. Whenever moment the bend of the periphery is \u03c0D/p for all the shafts, where D is that the inner distance across of the stator. Figure 9 shows the attractive field conveyance of IM under the static unusual condition with a seriousness level of 43%. It tends to be seen from Figure 9(a), asymmetry happens in the attractive field appropriation because of unpredictability flaw, the attractive tomahawks position of the posts become temperamental and each shaft length changes around \u03c0D/p when arriving at the shortcoming district of the rotor during the pivot. The length of the attractive posts of the flawed machine experiences an intermittent variety round the rotor during motion turn. From the conveyance of attractive fixation in Figure 9(b) it is seen that the motion thickness becomes uneven bringing about immersion inside the rotor and stator teeth close to the dislodged position because of issue. At the point when the issue seriousness expands, the asymmetry in transition lines and motion thickness dispersion over IM increments. ICMSMT 2020 IOP Conf. Series: Materials Science and Engineering 872 (2020) 012049 IOP Publishing doi:10.1088/1757-899X/872/1/012049 Thus, electromagnetic qualities of PWM inverter fed IM is broke down for solid and eccentricity flaw of degree 43% utilizing FEM model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002477_etcm.2016.7750839-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002477_etcm.2016.7750839-Figure2-1.png",
        "caption": "Fig. 2. Quadcopter formation diagram.",
        "texts": [
            "00 \u00a92016 IEEE For the reduced kinematic model [12], a technique called \u201csmall angle approximation\u201d is used, which is a very convenient technique to simplify trigonometric laws and presents an acceptable accuracy when the angle approaches to zero, this model assumes that \u2192 0 and \u2192 0, therefore, \u2245 1, \u2245 1 and \u2245 0, \u2245 0, given the next equation as follows: = \u2212 000 0 1 (2) III. CONTROLLER DESIGN To implement the quadcopter formation control law, the reduced kinematic model is used. The aim is to maintain the three quadcopters in a triangle formation where each one is located within its corresponding vertex relative to the center (centroid), as shown in Fig. 2. In addition, the assembly should be moved, keeping the desired shape and posture along the implemented trajectory, ([1], [2]) Where = ( , , ) , = ( , , ) and =( , ,, ). The mathematical model that is going to be used for each quadcopter is shown in (3), which is: = \u2212 000 0 1 = (3) Where = [ ] are the velocities of the quadcopter, is the rotation matrix of the quadcopter and is the orientation of the quadcopter, which also can be expressed as follows: = (4) From Fig.2, it can be observed the posture of each quadcopter , and , the distance between and is , the distance between y is and the distance between y is , is the angle between and , , are the centroid postures and is the formations posture angle. stays constant and it isn\u2019t use during the analysis. Therefore, shape variables are defined by: \u210e ( ) = [ ] (5) And posture variables are defined by: \u210e ( ) = [ ] (6) Mathematically, the shape variables and are found by analytic geometry and angle by cosine law: \u210e ( ) = = ( \u2212 ) + ( \u2212 )( \u2212 , ) + ( \u2212 ) (7) And posture variables are using the concept of centroid as: \u210e ( ) = = ( + + )/3( + + )/3\u2044 \u2044 ( )\u2044 \u2044 ( ) (8) The relationship between shape variables \u210e with the postures of each quadcopter and by corresponding differential relationship is: \u210e ( ) = (9) Where \u210e = [ ] are the time variations of shape variables, is the Jacobian matrix so that relates the speed of quadcopters with temporal variations of shape variables and = [ ] is the vector of velocities relative to the fixed quadcopter grounding system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002160_icra.2016.7487445-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002160_icra.2016.7487445-Figure1-1.png",
        "caption": "Fig. 1. Three-dimensional forces acting on a wheel with grousers.",
        "texts": [
            " In this paper, we first explain the method of measuring the three-dimensional forces acting on the grouser and the wheel surface between each grouser of the wheel, and present the measurement results as the raw force data. Then, we propose a calculation method of the stress distributions generated beneath the wheel with grousers, based on the measured 978-1-4673-8026-3/16/$31.00 \u00a92016 IEEE 2828 force data at a certain angle of the wheel rotation, and represent the calculation results of the stress distribution of the wheel with grousers. Fig. 1 shows the measurement locations of the forces. The upper left diagram shows the three-dimensional forces acting on middle of the wheel surface. The upper right diagram shows the three-dimensional forces acting on the grouser. The lower left diagram shows the three-dimensional forces acting on rear portion of the wheel surface. The lower right diagram shows the three-dimensional forces acting on front portion of the wheel surface. The forces acting on the wheel with grousers can be divided into the force acting on the grouser and the force acting on the wheel surface between each grouser"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure9-1.png",
        "caption": "Fig. 9: Tube-to-tube welding case",
        "texts": [],
        "surrounding_texts": [
            "In this section simulation results in dealing with the above three typical welding cases will be given to verify the proposed cooperative motion planning method. All simulations are programmed in our developed off-line programming software ROBOLP, and run on a PC with Intel (R) Core (TM) i3-2120 CPU@3.30GHz 3.30GHz and 4G RAM. The distance from the end of tool to the seam is set to 10 mm, and the welding speed is set to 10 mm/s."
        ]
    },
    {
        "image_filename": "designv11_30_0000382_ijhvs.2018.089897-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000382_ijhvs.2018.089897-Figure5-1.png",
        "caption": "Figure 5 Vector along the direction of the screws axis (Models of the front and rear of the trailer)",
        "texts": [],
        "surrounding_texts": [
            "Screw theory enables the representation of the instantaneous position of the mechanism in a coordinate system (successive screw displacement method) and the representation of the forces and moments (wrenches), replacing the traditional vector representation."
        ]
    },
    {
        "image_filename": "designv11_30_0000407_978-3-319-95966-5_4-Figure4.9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000407_978-3-319-95966-5_4-Figure4.9-1.png",
        "caption": "Fig. 4.9 Serrated chip formation in machining of Inconel 718 (with permission to reuse) [55]",
        "texts": [
            "5 m/min, the chip type is still a continuous ribbon type but the inhomogeneity increases as can be observed from the rise in serration along the free surface of the chip; see Fig. 4.8b. When the cutting speed approaches 61 m/min, the chip shows transition from continuous type to shear localized serrated type. Extensive twinning and shear localized band are clearly seen in Fig. 4.8c. Shear localization is further intensified beyond the transition speed in the narrow bands between the segments; see Fig. 4.8d. The mechanism of serrated chip formation in the machining of superalloys, in this case Inconel 718, can be conceptualized based on Fig. 4.9 as follows [55]: \u2022 When the tool proceeds toward the workpiece, the portion of the material located in zone 1 (PA1A2), near the tool tip, is exposed to high compressive stresses. \u2022 As cutting tool continues to move forward, this portion is forced upward over the inclined plane A1A2. Hence, the material between zone 1 and 2 has a tendency to escape to the free side of the chip and gets pushed toward B1D2. \u2022 Concurrently, shearing of workpiece material occurs at an angle /1 alongside the main shear plane A1B1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001776_j.ifacol.2015.06.362-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001776_j.ifacol.2015.06.362-Figure7-1.png",
        "caption": "Fig. 7. The modeled blade based on the proposed method with constant rake and relief angle.",
        "texts": [
            " (13) The obtained Vl is transformed back to the point on the cutting edge using MT3-T1=( MT1-T3) -1 . Based on the discussion in section 4, if gradients of working rake and relief angles decrease, tool wear characteristic would be increased. In the proposed method, it is tried to make the gradients equal to zero by setting constant values to the working relief and relief angles along the cutting edge. In this paper for the modeled blade, the working rake and relief angle are set to 5 and 31.2 degrees, respectively. Fig. 7 shows the modeled blade based on the proposed method. It should be noted that the proposed method is applied to both main and secondary cutting edge to show that how the model would be for full profile blades too. In case of full profile blades, the both sides of the blade should be considered and optimized because the both sides engage with the workpiece. The half profile blade would be simpler than what it is showed in the figure since only half of the blade needs to be changed. The free form shape of interior parts of the rake face is a result of applying tangency on all rake surfaces at their boundaries. In addition, the relief surfaces are designed to follow the epicycloid and do not have interference (gouging) with the cutting surface (the machined surface). 7. VALIDATION The blade shown in Fig. 7 is designed using 5 o working rake angle and 31.2 o working relief angle. Therefore, based on the discussion on Section 5, the modeled blade has a better wear characteristic especially at the corners. In addition, in order to demonstrate the improvement of the tool wear characteristics of the designed blade, two FEA simulations on Third Wave AdvantEdge are carried out. Tool wear is a result of thermal and mechanical loads on the tool during chip formation (Brecher et al., 2013a), therefore, in order to analyze tool wear for both conventional and proposed blade, tool load should be investigated and compared",
            " (13) The obtained Vl is transformed back to the point on the cutting edge using MT3-T1=( MT1-T3) -1 . Based on the discussion in section 4, if gradients of working rake and relief angles decrease, tool wear characteristic would be increased. In the proposed method, it is tried to make the gradients equal to zero by setting constant values to the working relief and relief angles along the cutting edge. In this paper for the modeled blade, the working rake and relief angle are set to 5 and 31.2 degrees, respectively. Fig. 7 shows the modeled blade based on the proposed method. It should be noted that the proposed method is applied to both main and secondary cutting edge to show that how the model would be for full profile blades too. In case of full profile blades, the both sides of the blade should be considered and optimized because the both sides engage with the workpiece. The half profile blade would be simpler than what it is showed in the figure since only half of the blade needs to be changed. The free form shape of interior parts of the rake face is a result of applying tangency on all rake surfaces at their boundaries. In addition, the relief surfaces are designed to follow the epicycloid and do not have interference (gouging) with the cutting surface (the machined surface). 7. VALIDATION The blade shown in Fig. 7 is designed using 5 o working rake angle and 31.2 o working relief angle. Therefore, based on the discussion on Section 5, the modeled blade has a better wear characteristic especially at the corners. In addition, in order to demonstrate the improvement of the tool wear characteristics of the designed blade, two FEA simulations on Third Wave AdvantEdge are carried out. Tool wear is a result of thermal and mechanical loads on the tool during chip formation (Brecher et al., 2013a), therefore, in order to analyze tool wear for both conventional and proposed blade, tool load should be investigated and compared. The tool load can be investigated as temperature and stress distribution along the cutting edge (Brecher et al., 2013a). In face-milling or face-hobbing, uneven load distribution along the cutting edge occurs because of uneven undeformed chip thickness and the shape of the cutting edge and rake and relief surfaces. Since in this paper, only the gradients of the working rake and relief angles are considered, in order to nullify the effect Fig. 7. The modeled blade based on the proposed method with constant rake and relief angle. of uneven chip thickness, the simulations are carried out with uniform undeformed chip thickness (.12 mm) to see only the influence of constant angles. Cutting speed is 300 m/min and workpiece material is Al6061-T6. Fig. 8 shows the first simulation. The left and the right side of the Fig. 8 simulate machining process using conventional and proposed blade model, respectively. Since the corner of the blade is more prone to be worn out, the blade corner and small portions of the cutting edge connecting to the corner are considered in the simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003428_j.matpr.2020.08.109-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003428_j.matpr.2020.08.109-Figure1-1.png",
        "caption": "Fig. 1. Granule and the hole edges contact design scheme.",
        "texts": [
            " uipment vibra- The first task arises when it is sufficient to remove burrs and round the work pieces sharp edges, while the grooves and holes internal surfaces do not require machining according to the technical requirements. This is the most characteristic problem being solved while performing vibration finishing parts processing, which in general can be solved by using granules with the size RGR exceeding the size of the largest hole or groove L, i.e. RGR > L , to avoid jamming in them. The processing environment contact will be considered as a single granule contact until the burr is completely removed and the edge rounding required value is obtained (Fig. 1). Given the abrasive tool wear coefficient KW which largely determines the working medium flow rate, the of the product processed surface quality and the process productivity, the dependence for determining the processing medium granules optimal radius will be the following: RGR \u00bc KW \u00f0L\u00fe 2hb\u00de2 \u00fe h2 C 8hC \u00f01\u00de In the ratio above hb means the burr thickness at the base, mm; hc means chamfer height specified by technical requirements, mm; L is the linear or diametrical size of the hole being machined, mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002643_robio.2016.7866657-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002643_robio.2016.7866657-Figure5-1.png",
        "caption": "Fig. 5. Example of a spatula and nodes displacement in [1] (left). 3D Model of the convex rigid body (middle). The two collidable bodies of the clamp (right).",
        "texts": [
            " Finally, in order to update the soft object shapes, the elastic force f = kt (2) is applied to all the nodes of the soft object that are involved in the collision. In Eq. (2), t is the distance computed with the ray casting procedure. This force causes a deformation of the soft object in order to bring it outside the volume of the rigid body. The rigid body used in [1] is a parallelepiped and the nodes used for collision detection are placed only on the two larger surfaces. The symmetrical shape of the body allows an easy manual placement of the nodes by defining the resolution along the planar directions of the two surfaces, see Fig. 5 (left). In this work, the proposed algorithm has been extended from rigid boxes to generic convex objects modelled with triangular meshes. The algorithm in [1] requires that nodes on parallel surfaces are associated in pairs that belong to the same surface normal. For convex objects, as in Fig. 5 (middle), the nodes association is realized by placing a node in the centroid of each triangle mesh and by associating it to a node that belongs to the triangle intersected by the ray casting. The ray source is the node itself and the direction is the opposite of the normal to the triangle to which the nodes belongs. In this way, each triangle finds its opposite pair. With respect to [1], where if a node x is associated to a node y also the vice-versa holds, in this case the nodes are not always mutually associated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001045_amr.1051.139-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001045_amr.1051.139-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of Double shear test Specimen",
        "texts": [
            "173, Monash University Library, Clayton, Australia-09/12/14,00:56:53) laid within the period of the resin-hardener mixture drying.The resin-hardener mixture is again applied over flax fibre layer and a layer of Kenaf fibre is placed over it. Finally a layer of flax and GFRP is added. The complete structure is compressed by applying a uniform load of 10 kg until the composite sets. The arrangement should be such that the GFRP always forms the top surface. The arrangement of the hybrid composite can be represented using schematic diagram shown in Figure1. Then according to the requirements the composite is cut for testing and applications. Similarly GFRP + Flax and GFRP + Kenaf laminates are fabricated by replacing the layer of allied material in the composite with the layer of same natural fibre. In general the natural fiber obtained from plants and animals has various advantages like biodegradability, vast availability and economical over synthetic fibers. They are extracted by some means mechanical processes and then treated chemically before use",
            " Epoxy resin consisting of at least two epoxide groups may be either pre-polymers or polymers. HY951 hardener also known as araldite has high mechanical and good chemical properties. The mixture of both forms the prime constituent that binds various layers of fiber. The mechanical test namely double shear test is performed for the fabricated composites Double shear test is also done in universal testing machine using a special fixture. ASTM: D5379 standards is adapted here and sample is prepared as in Figure 1. The specimen is placed in the fixture and load is applied until the specimen breaks. The breaking load is noted and Load Vs Deflection graph is generated. The results of double shear test are shown in table 1. The graph obtained for Load Vs Displacement is shown in figure 2. The displacement increases linearly as the load increases, and then there is a small decrease in displacement with respect to load. This small dip in graph corresponds to yield load. The break load for the hybrid composite is little bit higher than the other two"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002501_aus.2016.7748031-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002501_aus.2016.7748031-Figure2-1.png",
        "caption": "Fig. 2 EHA of WEBER-HYDRAULIK",
        "texts": [
            " Direct drive volume control EHA developed by this company has been used in precision forging machine, ship rudder, continuous casting equipment, printing machine, six degree of freedom platform, 2500 tons hydraulic molding machine etc. [4]. Fig.1 DDVC of Japan\u2019s Electric Co., Ltd. The EHA product of Germany\u2019s WEBER-HYDRAULIK company\uff0c has the characteristics of low energy consumption, high energy efficiency, easiness to adjust power level, resistance to external influences (wind, vibration, etc.), overload protection valve, damping effects at the end position, no hysteresis, UL and CE certified, etc., as shown in Fig. 2. And some other features can also be chosen, such as power adjustment, electrical controller, valve control for auxiliary functions. The selectable input power is 3kw or 20kW and flow range from 0.2 L/min to 200L/min, operating temperature range from -30\u2103 to 80\u2103 , the maximum cylinder diameter is up to 250mm[5]. The EHA of German\u2019s VOITH company, which is named SelCon\uff0cis shown in Fig.3. The core components consist of a servo motor, a Voith internal gear pump, a cylinder with an internal return spring and a displacement transducer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000020_b978-0-12-816805-9.00008-9-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000020_b978-0-12-816805-9.00008-9-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of (A) the experimental setup for piezoelectric output current measurement and (B) sample and electrode design [11]. (Reproduced with permission from SAGE Publishing (https:// www.stm-assoc.org/intellectual-property/permissions/permissions-guidelines/). Copyright 2018.)",
        "texts": [
            " Various experimental characterizations have proven that 3D printed BTO/PVDF nanocomposite films are outperforming when compared with solventbased films. Scanning electron microscopy (SEM) results have shown that 3D printed films have less agglomeration, porosity, and cracks as compared to solvent-based films. Similarly, XRD (X-ray diffraction) spectra display unified distribution of BTO nanoparticles, further Fourier transform infrared spectroscopy (FTIR) results shown \u03b2-phase content are more in 3D printed films. Current generated in this process is also more. Three-dimensional printed sample test piece and test setup [11] are shown in Fig. 1. The growing demand of next-generation multifunctional materials has attracted the attention of researchers from academia and industry due to its various applications such as energy storage, flexible electronics, etc. [14]. The significant development of new functional materials is the basis for implementing the innovative additive manufacturing techniques. Generally, the additive manufacturing processes such as fused deposition modeling (FDM), stereolithography apparatus (SLA), liquid deposition modeling (LDM), selective laser sintering (SLS), and electron beam manufacturing (EBM) are used to construct 3D structures by adding the material, layer by layer, whereas in traditional fabrication techniques subtractive approaches are used to remove the material from an original work piece for obtaining the final shape [15]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003511_j.triboint.2020.106696-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003511_j.triboint.2020.106696-Figure13-1.png",
        "caption": "Fig. 13. The photograph of labyrinth seal specimen.",
        "texts": [
            " The thickness of coating 0.15 mm, and the surface roughness of the coating is 0.02 \u03bcm. The crucial initial geometric tolerances of the brush seal and rotor directly decided the initial interference and the accuracy of the test, which have been measured before testing. The results are shown in Table 3. We use a rotor in three cases (E1, E2, E3), but a metal ring can be used to adjust the axial position of the brush seal. The brush seal contacted with the rotor at three different positions. A labyrinth seal shown in Fig. 13 was used to match the brush seal. The labyrinth seal is made of 38CrMoAl, a kind of structural steel, and the surface roughness is 0.02 \u03bcm. At least one experiment is necessary to determine the deformation correction factor \u03c8. To improve the reliability of the wear model, Two experiments were used to determine \u03c8 and one to validate the mathematic model, therefore, three experiments have been conducted in different situations. All the experiments were conducted at room temperature. The other operating conditions of the three experiments are shown in Table 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000007_978-981-15-0434-1_3-Figure1.2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000007_978-981-15-0434-1_3-Figure1.2-1.png",
        "caption": "Fig. 1.2 Schematic representation of friction",
        "texts": [
            "), Automotive Tribology, Energy, Environment, and Sustainability, https://doi.org/10.1007/978-981-15-0434-1_1 3 and wear at interface, a lubricant is supplied. Tribology involves the study and the applications of the principles of friction, wear, lubrication and surface modification which is diagrammatically shown in Fig. 1.1. The resistance force experienced in relative motion of solid surfaces, fluid layers, and/or material elements sliding against each other is known as friction which is shown in Fig. 1.2. There are two types of friction, dry friction and fluid film friction (Rabinowicz 1995). When there are two solid surfaces sliding against each other and energy is entirely dissipated between the surfaces then it is known as dry friction. This dry friction is again subdivided into two parts; static friction and kinetic friction. The degradation of material or removal of material is known as wear. Wear occurs due to the friction between two surfaces. There are different types of wear which is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001287_icuas.2014.6842244-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001287_icuas.2014.6842244-Figure2-1.png",
        "caption": "Fig. 2: Spring-damper flocking model, Determination of the closest point (bottom)",
        "texts": [
            " Then, a single robot will interact only with the directly surrounding robots. The ideal distance Di represents the target distance for robots surrounding another one avoiding collision and maintaining an acceptable flock. In our model, this distance corresponds to the equilibrium length of the spring-damper system. If two robots are too close to each other, the spring compression generates a force to separate them proportionally to the compressed length. Also, if two robots are too far from each other, the system will tend to reduce the distance in between. The figure 2(a) represents this interaction between two robots. To compute this system, since the force generated by the spring is proportional to the error distance regarding to the equilibrium length, this system can be modeled by a proportional corrector. Similarly, the effect of the damper is proportional to the velocity of the moving robot. The spring-damper system is then equivalent to a proportionalderivative (PD) controller on robots positions. For two robots, the setpoint for this controller is the closest point from one robot which is at the Di distance from the other robot (Point P on figure 2(a)). To compute the controller, we calculate the distance between one robot to this point P. Considering two robots A and B separated by a distance D (figure 2(c)), we can find the coordinates of the point P. From this, we can compute a position PD controller to have a control loop feedback which tends to maintain robots at the desirable distance Di. Moreover, to eliminate steadystate errors and compensate perturbations, we introduce a third term, proportional to the integral of the position error. Finally, to maintain robots at a desired distance, we use a classic proportional-integrate-derivative (PID) controller. Considering a robot and a desired position we define by e the position error",
            " We can use this control law to compute the robot acceleration: x\u0308 = K p\u2217 e+Kd \u2217 e\u0307+Ki\u2217 \u222b e For the usual case of a robot interacting with more than one other, this controller provides an easy way to compute a new control law. For one single robot, we calculate the desired position regarding the position of each robot surrounding it. Then, the desired position we use for the setpoint of the position controller is the barycenter of all the positions calculated for each couple of robots. This is shown in the figure 2(b). In this example, to compute the setpoint position of robot A: we calculate first the desired position considering only interaction with robot B. This gives the point P1. The same principle for robot C gives the point P2. We compute then the barycenter of P1 and P2, which gives the point P, the setpoint position for the PID controller of robot A. The same is done for robots B and C. The robots will organize themselves into a minimum energy configuration, regarding the energy stored in the mechanic spring-damper system",
            " Quadrotor robots are implemented with a realistic dynamic model, controllers are modeled taking into account real-time constraints like execution time or scheduling policy and networks deal with MAC and routing protocols or unreliability like time delays or message losses. That allows accurate robotics, control and communication studies of the system. A realistic model of quadrotor is then used to simulate one robot behavior. As in [6], dynamic and kinematic models have been implemented, taking into account mechanics and aerodynamics constraints. We model the quadrotor with four rotors in a cross configuration. One rotor is considered as the front of the robot, and three axis are then defined as shown in figure 2(d): the X axis from back to front, Z axis vertically and Y along the rotor axis perpendicular to X axis. Four variables can be tuned to control the robot. The throttle U , leads to a vertical force to set the height of the quadrotor. The throttle is controlled by an equal variation of the four propeller speeds. The three other commands roll (\u03a6), pitch (\u0398) and yaw(\u03a8) are respectively related to rotation angles about the X, Y, and Z axis. We consider \u21261, \u21262, \u21263 and \u21264 respectively the front, right, rear and left propeller speeds, and Ixx, Iyy and Izz the moments of inertia around the X,Y and Z axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure1-1.png",
        "caption": "Fig. 1. Overall design of the PIBOT",
        "texts": [
            " And the overall design layout can be summarized to state that the complete system is expected to weigh around 210 kg and measure about 3.5 m in length. The main purpose of PIBOT system is to find a corrosion defects on the outer surface of pipe and collect video images inside of pipelines. And by the wireless communication system with 5.1 GHz bandwidth, the collected data is sent to the control station above ground. The overall design of the PIBOT and major functions of each module are depicted in Fig. 1. The PIBOT consists of the 8 modules which are camera modules, drive modules, support modules, pump module, battery modules and MFL module. Each module is connected through an articulated joint. The role of camera module is to collect video image inside the pipeline. Also it has a sonde system to locate a point where the PIBOT system runs in the buried pipeline. The camera module was designed to integrate computing, wireless communication, video-sensing, lighting and emergency-locator systems into a single monolithic module"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000345_978-94-007-7194-9_2-1-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000345_978-94-007-7194-9_2-1-Figure4-1.png",
        "caption": "Fig. 4 Implementation of the singularity-consistent inverse kinematic approach: three motion patterns at the shoulder singularity of a nonredundant limb. Left (a), motion through the singular path; middle (b), motion along the singular path; right (c), rotation around the singular path [28]",
        "texts": [
            " a rotating coordinate frame, the z axis being always aligned with the singular (constrained) direction. Interested readers are referred to [27] where an implementation with a kinematically redundant limb is also explained. At the (avoidable) shoulder singularity of a nonredundant limb, on the other hand, the SC method can be applied to control instantaneous motions resulting in three motion patterns [28]. These are obtained with the first three joints (the positioning subchain) of a 6R limb, as shown in Fig. 4. The dashed (red) line represents the singular path at the shoulder singularity. Whenever the end-link is on that path, motion is constrained in the direction parallel to the arm-plane normal. Motions along the unconstrained directions are shown in Fig. 4a, b. In Fig. 4a, the end-link moves through the singularity in the direction transverse to the singular path. In Fig. 4b, the end-link moves in the direction parallel to the singular path, whereby the singularity is not escaped. Note that the end-link may reach the workspace boundary along the singular path, thus arriving at a double shoulder-elbow singular configuration. Finally, Fig. 4c displays a self-motion (cf. Sect. 4.1) pattern resulting from a commanded end-link velocity having a nonzero component in the constrained direction (parallel to the arm-plane normal). The joint at the shoulder base rotates the limb and hence the arm plane, until the commanded end-link velocity component along the constrained direction is nullified. Thereafter, the end-link will leave the singularity along the unconstrained directions. From differential kinematic relation (3), it is apparent that the ability of the endlink to move instantaneously along a given spatial (rigid-body motion) direction will depend on the current limb configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000734_j.compstruc.2013.05.008-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000734_j.compstruc.2013.05.008-Figure1-1.png",
        "caption": "Fig. 1. (a) Description of the ring rolling process scheme. (b) Mesh refinement for the F.E. simulation and calculated reduced contact area.",
        "texts": [
            " The presented work is focused on this key factor that influences algorithmic robustness and accuracy. It aims at improving contact algorithms by smoothing without increasing the computational cost, which means without decreasing the mesh size or the time step. This study is introduced by a short study of a ring rolling process which properly shows the contact issue; it is then extended to a more general frame and to other problems. Ring rolling is a hot forming process that produces rings with varied size features. It is schematically shown in Fig. 1a. The initial ring is placed against the mandrel roll, which is forced toward the drive roll. The drive roll rotates continuously, reducing the wall thickness, and increasing the diameter. As shown in Fig. 1, the key challenge for this process simulation is that the material deformation occurs in a much reduced part of the computational domain. Preliminary calculations with Forge software [1] have provided non homogenous and abnormally high strain values \u00f0 e P 5\u00de (see Fig. 2), which appeared to be closely related to the contact treatment. Given the relative coarseness of the mesh (see Fig. 1b), the contact area is limited to very few nodes, so contact is not detected at the beginning of the process and temporal oscillations are observed. This is likely regarded to come from the coarse and non-smooth description of the cylindrical tools discretized by linear C0 triangular facets. In order to check this hypothesis, an analytical description of the tools is introduced. It is easily applicable to such a process where obstacles are cylinders and cones. Simulation results (see Fig. 3) show that the quality is thus significantly ameliorated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002312_1464419316661969-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002312_1464419316661969-Figure1-1.png",
        "caption": "Figure 1. Interaction between the ball and raceway.",
        "texts": [
            " Bearings are assumed to operate under isothermal conditions, with all thermal effects ignored. However, under normal operations, heat is generated in the bearings due to shearing of a lubricant film in roller\u2013raceways contacts, which has been studied by Liu et al.24 5. Bearing components are also rigid, and outer raceway is rigid to the housing. These elements are elastic and can be subjected to deformation actually, which would affect the results of analysis. Such solutions were provided by Gao et al.25 Interactions between ball and raceways. As shown in Figure 1, geometrical interactions between the ball and raceways are described for obtaining forces and moments. The necessary frames are built: inertial frame OXYZ is the reference frame, race-fixed coordinate frame RXrYrZr is used to confirm the raceway centre in the inertial frame, ball azimuth frame AXaYaZa is also established for determining positions of the ball centre in the inertial frame, and descriptions of the frames above are given in the study by Gupta.26 When all the frames are confirmed, two vectors rib and rir are expressed as position vectors of the ball centre B and raceway centre R in the inertial frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002964_tia.2020.2986181-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002964_tia.2020.2986181-Figure7-1.png",
        "caption": "Fig. 7: Linear TH simulation solution, imposing the same stator current and frequency as in the non-linear problem. Because of the marginal role of the magnetizing current in the LR regime, the rotor bars induced current is similar to that in the non-linear solution, as well as the current density profile.",
        "texts": [
            " The basic idea is to start from the TH simulation, to get the actual rotor bars current distribution, then store how the induced current is divided in the several sectors with respect the total bar current and finally reproduce the same distribution in the MS simulation. However, getting the bars current distribution from the non-linear TH analyses is rather expensive in terms of computational time. At first, it worth to verify that, from the linear TH simulation, a similar results to the non-linear TH can be obtained, about the bar induced current distribution. The field and induced current solution of the linear problem is reported in Fig. 7. In order to demonstrate that the linear solution can be used instead the non-linear one to determine the bar current distribution, it has to be verified that, from the two field solutions, the ratio between the slot sector current and the total bar current is similar, for each slot sector and each slot. In other words, for each sector of the slots, a normalized current density is derived as: jp = < ( ip ) < (ibar) 1 Sbar [1/m2] (8) The comparison between the normalized current in each slots sectors in the linear and non-linear solution is shown in Authorized licensed use limited to: Carleton University"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000420_math.mag.90.2.87-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000420_math.mag.90.2.87-Figure11-1.png",
        "caption": "Figure 11 Pi fits inside Po.",
        "texts": [
            " Table 3 shows the coordinates of the projected vertices k \u2032 in terms of \u03d5, found from the familiar formulas for the rotation through the angle \u03d5 about the origin in the xz-plane. The coordinates are presented in the (x, y) axes in Po, and to the same scale as the vertices given in Table 2. Choosing \u03d5 so that the projection of the edge 1-5 is inside and parallel to the edge 7-4 minimizes the maximum distance between these two segments, and a calculation shows that this is accomplished for \u03d5 = 3.1289\u25e6 (and similarly for 7-10, 3-6, and 12- 8), while leaving the image of vertices 2 and 11 inside Po (Figure 11a). The interior of Po is described by the system of inequalities (4). \u23a7\u23aa\u23a8 \u23aa\u23a9 \u22122 \u221a 3(1 + 3\u221a 5 ) \u2212 \u221a 3x < y < 2 \u221a 3(1 + 3\u221a 5 ) \u2212 \u221a 3x \u2212\u221a 3(1 + 3\u221a 5 ) < y < \u221a 3(1 + 3\u221a 5 ) \u22122 \u221a 3(1 + 3\u221a 5 ) \u2212 \u221a 3x < y < 2 \u221a 3(1 + 3\u221a 5 ) + \u221a 3x (4) Verifying that these coordinates satisfy the constraints (4) establishes that the vertices of Ii are in the interior of P0, and this completes the coordinate proof that the icosa- This content downloaded from 128.122.230.148 on Thu, 09 Mar 2017 16:20:00 UTC All use subject to http://about"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001414_s00170-013-5010-1-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001414_s00170-013-5010-1-Figure4-1.png",
        "caption": "Fig. 4 The path of the contact point between the driving pin and the lathe fork in the axial intercept plane of the spindle",
        "texts": [
            "hu Power dissipation in the gear can be reduced significantly with these modern drive pairs which are characterized by favorable hydrodynamic conditions, great strength, and high efficiency [4]. In case of power loss, it is important to apply those geometrical characteristics of the cog which result in good connection terms. During our examination, we started from the modeling of the driving of cylindrical driving pin (Fig. 3). We supposed that there is a punctual connection between the driving pin and the lathe fork. That is how we could determine the path of the contact point between the lathe fork and the driving pin during one turning. Figure 4 shows that in the beginning of turning, on the axial intercept plane of the spindle the contact point between the driving pin [I.] and the lathe fork [III.] is point A\u2032. The lathe fork [III.] is situated perpendicular to the worm axis [IV.], namely it encloses half cone angle \u03b41 with the head surface of the spindle [II.]. After turning the spindle with angle \u03c6p, the contact point is taken from point A\u2032 to point P\u2032 and during this movement it completes path sp. After turning of \u03c6p=180\u00b0, the contact point is taken from point A\u2032 to point B\u2032 and during this movement it completes the longest path \u0394s. It can be seen if the angular rotation \u03c6 is ranging between 0\u00b0 and 180\u00b0 then the path of the contact point is increasing; in case of 180\u00b0 the longest path \u0394s is completed; between 180\u00b0 and 360\u00b0 the path is decreasing and in case of 360\u00b0 the contact point is taken back again to the initial distance x, i.e., the starting position (point A). Based on the CDE rectangular triangle (Fig. 4): y \u00bc a sin d1 \u00f01\u00de Based on the CAA\u2032 rectangular triangle (Fig. 4): x \u00bc y rm\u00f0 \u00de tgd1 \u00f02\u00de Substituting (1) into (2): x \u00bc a sin d1 rm tgd1 \u00f03\u00de Based on the CBB\u2032 rectangular triangle, the total path of the contact point of the lathe fork and driving pin is (Fig. 4): x\u00fe \u0394 s \u00bc y\u00fe rm\u00f0 \u00de tgd1 \u00f04\u00de \u0394s \u00bc y\u00fe rm\u00f0 \u00de tgd1 x \u00f05\u00de 3.1 The position of the contact point on any arbitrary \u03c6p angle position We have developed a universal correlation. Based on this the position of the contact point for any arbitrary \u03c6p, angle position can be determined. Based on the PFD rectangular triangle (Fig. 4): rP \u00bc rm cos 8P \u00f06\u00de Based on the CPP\u2032 rectangular triangle (Fig. 4): sP \u00bc y rP\u00f0 \u00de tgd1 x \u00f07\u00de By substituting (6) into (7), the position of the contact point in case of any arbitrary \u03c6p angle position is: sP \u00bc y rm cos 8P\u00bd tgd1 x \u00f08\u00de As a result of the shifting of the worm shaft by half cone angle, the path curve of the driving pin will be an ellipse path instead of a circle on the perpendicular plane to axis. The half cone angle of conical worms is between 5\u00b0 and 10\u00b0 [8, 9], that is why the major axis of ellipse on the perpendicular to worm axis section is larger than the pitch circle radius of the spindle but only to a small extent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000726_icciautom.2013.6912808-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000726_icciautom.2013.6912808-Figure3-1.png",
        "caption": "Figure 3. Microsatellite reference and body coordinates [12]",
        "texts": [
            " Here, the parameter \u03b8 is adjusted by minimizing a loss function as: (3) (4) (5) (6) (7) (8) (10) (9) (11) (12) (13) 2 2 1)( eJ =\u03b8 To make J small, it is reasonable to change the parameters in the direction of the negative gradient of J as: \u03b8 \u03b3 \u03b8 \u03b3\u03b8 \u2202 \u2202\u2212= \u2202 \u2202\u2212= eeJ dt d Fig. 2 shows the block diagram of the RMPC method based on MRAS algorithm. III. THREE DEGREE OF FREEDOM SATELLITE STATE SPACE MDEL In this section, three degree of freedom rigid satellite model is presented. A microsatellite is shown in Fig.3 [12]. Axes XB, YB and ZB are satellite\u2019s body axis frame. The angles roll )(\u03d5 , pitch )(\u03b8 and yaw )(\u03c8 are defined by successive rotations around the body axes XB, YB and ZB. Parameters p, q and r are the angular rate. The nonlinear state model of the satellite is according to the following relation [12]: + \u2212 ++ \u2212+ +\u2212 +\u2212 = \u03b8 \u03c6\u03c6 \u03c6\u03c6 \u03b8 \u03b8\u03c6\u03c6 \u03c8 \u03b8 \u03c6 cos cossin sincos cos sin)cossin( rq rq rqp I pqIqpIM I rpIprIM I qrIrqIM r q p zz yyxxz yy zzxxy xx zzzzx Where Mx, My and Mz are the input torques, Parameters yx II , and zI are the moment of inertia around the body axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000834_2014-01-2064-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000834_2014-01-2064-Figure1-1.png",
        "caption": "Figure 1. The driveline layout of a commercial utility vehicle",
        "texts": [
            " Secondly, the stiffness, inertia and damping parameters' sensitivities are analyzed based on 8 DOFs model which is simplified by the multi-DOF model. Thirdly, DOE analysis based on stiffness and damping is provided to calculate their contribution to the transmission torsional vibration response. Finally, a clutch with small stiffness and big rotational angular is tested in which the transmission shaft torsional vibration and rattle noise are significantly reduced. The analysis is based on a typical driveline configuration of a rear-wheel-drive commercial utility vehicle. Figure 1 illustrates the driveline layout. There are other factors which might affect the driveline torsional vibration, such as the engine mount, propeller shaft angle, and the phase of propeller shaft Yoke. In this paper, these factors are neglected. The RWD system in this paper consists of crank shaft, flywheel, transmission, drive shaft, differential, half shaft, wheel, etc. A coupling multi-body model including the lumped mass and distributed mass is established. The drive shaft and half shaft are discretized by finite element method to generate the modal neutral file (MNF) which is imported into the dynamic model that is connected with adjacent rigid components by kinematic pairs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002253_1464419316660930-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002253_1464419316660930-Figure5-1.png",
        "caption": "Figure 5. Kinds of REs and their parameters: (a) ball; (b) roller.",
        "texts": [
            " How to model bearings using different REs. . How to model bearings in different internal configurations. . How to model bearings with different numbers of rows. From ball bearings to roller bearings One of the main features of REBs is the kind of RE used. Thus, bearings are classified as ball bearings or roller bearings and, besides that, roller bearings can be composed of cylindrical, needle, tapered or spherical rollers. In this paper, two kinds of RE are defined when modelling, as shown in Figure 5. A sphere of diameter Dw is defined for ball bearings (see Figure 5(a)), whereas a general roller with mean at Univ of Newcastle upon Tyne on September 27, 2016pik.sagepub.comDownloaded from diameter Dw and length Lw is defined for roller bearings (see Figure 5(b)). Despite being quite different between them, the four kinds of rollers that can be found in off-the-shelf bearings can be easily modelled using an only roller model and varying its parameters, which are the tilt angle and the radius of curvature R, as shown in Figure 5(b). Thus, the combination of different values of and R leads to the four kinds of rollers, as can be seen in Table 3. It should be highlighted that the roller model only supports the combinations shown in Table 3 and other combinations of the attributes of the model are not taken into account due to lack of sense. Besides that, it should be noted that the relationship between the length and the mean diameter is the only difference that has been considered between the model of cylindrical rollers and the one of needle rollers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001172_j.tsf.2013.03.132-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001172_j.tsf.2013.03.132-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the whipping-based electrospinning platform.",
        "texts": [
            " The whipping behavior of electrospinning is utilized to generate filmwith various coiled, wavy, and zigzag fibers [27], which has a higher stretchability than film with straight fibers, and the voltage influence on fiber shape is investigated. Then detachment lithography is utilized to pattern the electrospun film. The electrospun film is transferred onto an elastomeric substrate (PDMS), and the film-PDMS system is pressed on the siliconmoldwith proper pressure and temperature, then the pattern is formedwhen the substrate is peeled off quickly. The stretchability of the patterned film is studied through experiment and finite element simulation. Fig. 1 is the schematic diagramof thewhipping-based electrospinning platformwhich consists of three parts, nozzle, substrate and high-voltage .doi.org/10.1016/j.tsf.2013.03.132 power supply. The high voltage stretches the jet flow from nozzle to substrate. The electrified jet emitted from the cone vertex is inherently unstable. After traveling a few centimeters in a straight line, it undergoes bending/whipping instability which is a special type of non-axisymmetric instability owing to the action of the electric field [13]. The bending stage is the root to produce curve fibers and it can be further divided into first bending stage and second bending stage, corresponding to the large helix and small wave respectively as schematically shown in Fig. 1. The three stages, which are stable stage, first bending stage and second bending stage, are able to generate straight fiber, large scale curve fiber and small scale curve fiber separately. The three stages can be adjusted by changing process parameters like nozzle-to-substrate distance, applied voltage, substrate speed etc. The second bending stage is adopted to generate film composed of small wavy fibers to get high stretchability. The experiments were conducted at room temperature with relatively humidity of about 48%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002592_978-3-319-50904-4_37-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002592_978-3-319-50904-4_37-Figure3-1.png",
        "caption": "Fig. 3. Detail view of the fabricated drive module",
        "texts": [
            " And it consists of battery pack to provide power for the camera module, the main control board, the sub control board, the locating-sonde coil, the forward LEDs, charging contacts and the fisheye camera-dome. The camera module is shown in Fig. 2. The drive module includes the ability to keep the robot body in the center of pipeline with the driving wheels which have the support arms. All the required mechanical elements, electrical PCBs, various sensors and odometers were integrated by the design shown in Fig. 3. The module was designed to expand/collapse a set of four drive-legs, and each wheel system contains a harmonic gear to rotate a drive-wheel using In-wheel motor. The battery module provides 25.2 VDC and up to 70 Ah. of energy to the robot for a meaningful 10 h mission. The battery module contains lithium polymer battery-cells combined into packs, safety electronics and voltage converters to allow monitoring of charge and discharge. An image of the layout of the module is shown in Fig. 4. The pump module was designed to provide the pneumatic actuator of the support system for the necessary centration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001390_cicare.2014.7007838-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001390_cicare.2014.7007838-Figure5-1.png",
        "caption": "Figure 5: (a) Bed sensor design; (b) architecture of the pressure sensor; (c) a photo of the bed sensor prototype",
        "texts": [
            "4,a) assesses if the door is opened by the PD. The tests confirmed that both tag and the RFID antenna are not perceptible as they walked. Also, using slippers as RFID-tag is an acceptable solution, since elderly people usually wear slippers at home all the time. The passive RFID tags do not require charging batteries minimizing the maintenance tasks. The bed sensor (BS) detects the presence of a person in bed. It consists of an air bag placed between the mattress and the base in the patient\u2019s bed, as shown in Fig.5 (a). The air bag is connected by a hose to an air pressure sensor that converts pressure changes into electrical signal. Fig.5 (b) depicts the operation of the pressure sensor. The pressure switch is connected to microcontroller and to XBee transmitter to signal the server when the pressure switch terminals open or close. When a person sits or lies on the bed, he or she compresses the airbag and increases the air pressure within it. The high pressure in the airbag pushes the plunger rightward against spring tension, closing the switch and causing the transmitter to send an \u201cON\u201d signal to the server, reporting the presence of the person on the bed. When the person gets off the bed, the air pressure within the airbag falls, the switch opens and the transmitter sends an \u201cOff\u201d signal, indicating that the person has left the bed. It matters very little where the person lies or sits on the bed. His weight is distributed over a large area of the air mattress, compressing it almost everywhere, and this raises the air pressure within it enough to close the pressure switch. Fig.5(c) shows a bed sensor prototype used in our tests. As the tests showed, the BS detects a person in bed correctly. In bathroom, corridor, or other places where the SC is difficult to install, we use simple motion sensors to detect patient motion. To communicate with the patient in the case of emergency, speakers, microphones and video cameras are used. We assume that at least one set of microphone, speaker and video camera is installed in each room of the home. The activity assessment module (AAM) receives information from sensors, assesses position and activity of the PD and alerts the caregiver if help is necessary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003130_physreva.101.063812-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003130_physreva.101.063812-Figure10-1.png",
        "caption": "FIG. 10. The stress and deformation in the x-z plane induced by one or two linearly polarized zeroth-order Bessel beams on a homogeneous sphere with physical parameters that are identical to those listed in the caption of Fig. 2 except with differing radii [indicated in the inset of panel (a)]. The first row [(a), (b)] shows the case of an x polarized beam coming from \u03b8 = 180\u25e6. The second row [(c), (d)] shows the case of one x polarized beam coming from \u03b8 = 180\u25e6 and one y polarized beam coming from \u03b8 = 0\u25e6. Here, for the single beam, three cases are shown: negative net force (pulling), zero net force, and positive net force (pushing). In all cases, each Bessel beam has a fixed power of P = 0.1 W and a wavelength of \u03bb = 0.532 \u03bcm.",
        "texts": [
            " As in the homogeneous case, the ranges of radii where optical pulling is possible are limited and become nonexistent at large radius. Finally, in Fig. 12, we examine one case of optical stress and deformation for the core-shell particle in a Bessel beam. The chosen radius corresponds to one of the regions on Fig. 11(a), where each Bessel beam pulls the core-shell 063812-13 particle. Along the axis of propagation, the broader stress profile on the shell leads to a stretching deformation unlike its homogeneous counterpart [Fig. 10(d)]. In addition, the optical stress will also deform the shell along the perpendicular axis. The resulting deformation combines these two effects and, due to the assumption of conservation of volume, this will lead to a compression of the shell along the axes where the stress is the lowest. We have presented a theoretical model for the optical stress and deformation induced by an arbitrary incident beam for a particle composed of concentric spherical shells. The electromagnetic component of the problem is treated in the framework of GLMT where the electromagnetic field is expanded in a basis of VSWFs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002808_0959651819899267-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002808_0959651819899267-Figure1-1.png",
        "caption": "Figure 1. (a) Photograph of the TRMS setup and (b) CAD model with coordinate frames.",
        "texts": [
            " Note that the sub-models developed in this article are generic and can be assembled with appropriate constraints to model various other systems such as helicopter, tricopter, quadcopter with full roll, pitch, yaw, heave, surge, and sway motions. This article is organized as follows: The next section describes the TRMS setup and various forces and couples experienced by it. Section \u2018\u2018Modeling of the TRMS dynamics\u2019\u2019 presents the modeling of the TRMS based on the BG approach. Section \u2018\u2018Results and discussion\u2019\u2019 discusses the results from the developed model and compares with the experimental test data. Finally, conclusions and perspectives are drawn in section \u2018\u2018Conclusion.\u2019\u2019 The TRMS comprises various components as shown in Figure 1(a). The shields, DC motors, and counterweight are rigidly attached to the supporting beam. The TRMS support beam is restrained in such a way that it can have only two independent rotational motions: one is rotational motion in the horizontal plane (yawing) and the other is rotational motion in the vertical plane (pitching). The global inertial coordinate system and the body-fixed frames used to model the setup are shown in Figure 1(b), where points m, t, c, and b are markers for assembly of various subsystems. Note that all body-fixed coordinate frames have origin at the respective center of mass or center of gravity (CG) and the coordinate axes are aligned along the principal axes. Any axis of symmetry of a body at the CG is a principal axis. The dynamics of the TRMS resembles a hovering helicopter. The purpose of the counterweight is to balance the pitch of the TRMS in steady state. Main and tail rotors can rotate about their own axes with respect to the support beam",
            " First, the DC motor circuit, interfacing, and aerodynamic propulsive forces acting on the rotors are modeled. Thereafter, the remaining support structures consisting of the beam, the counterweight, the rotors, and the rotor shields are developed using the concept of rigid body dynamics. The model variables are defined in the nomenclature given in Appendix 1. The main and tail DC motors are identical with different mechanical loads. Both the DC motor rotors are attached to the corresponding propellers whereas the stators (housings) are attached to the support beam ends, as shown in Figure 1. The active torques form the motors are applied to the propellers and the reactive torques act on the support beam through stators. These reactive torques cause cross-coupling between the yaw and pitch. The circuit diagram of the permanent magnet DC motor is shown in Figure 2 where parameters with subscripts m and t correspond to the main and the tail motors, respectively. The differential equations governing the electrical and mechanical domains of the permanent magnet DC motor are, respectively, given as Lam=t diam=t dt =Vm=t Kam=t vm=t Ram=t iam=t \u00f01\u00de Jm=t dvm=t dt =Kam=t iam=t TLm=t Cvm=t vm=t \u00f02\u00de where Vm/t is the input terminal voltage, Lam/t and Ram/t are the motor armature inductance and resistance, respectively, iam/t is the armature current, vm/t is the angular velocity of the rotor, Kam/t is the torque constant of the motor, Jm/t is the mass moment of inertia of the rotor about the axis of rotation, Cvm/t is the viscous damping, and TLm/t is the resisting torque due to the drag force",
            " Likewise, the thrust force on the tail rotor (Ft) is given by Ftn =Cttn 3 vtj j3 vt for vt \\ 0 \u00f08\u00de Ftp =Cttp 3 vtj j3 vt for vt \u00f8 0 \u00f09\u00de where Ftn and Ftp are the thrust forces on the tail rotor for clockwise and counterclockwise rotations, respectively, and Cttn and Cttp are two constant coefficients. The rotor blade drag forces are estimated by locking the support beam joints at its pivot and then applying constant input voltages in steps. Using the DC motor characteristics1 and equations (4) and (5), the drag coefficients in main and tail rotor blades are estimated at steady flow and steady blade rotation speeds (dv=dt=0). Nonlinear flat cable moment In the TRMS setup, a flat cable is attached to the TRMS hub, as shown in Figure 1. The cable opposes angular movement of the TRMS in the horizontal plane (yawing), as it applies a nonlinear moment on the hub. This nonlinear moment is difficult to estimate, as the cable behaves like torsional spring at some instants and rope at other instants. The cable\u2019s one end is attached to support beam through the hub and the other end passes through a hole in the support column, where the cable end can have slack movement. However, the moment due to cable can be modeled as a restoring force developed by a nonlinear torsional spring whose spring stiffness varies with the angle of twist/yaw (c)",
            " Similarly, CTF for transformation of velocities from the inertial frame to the body-fixed frame of reference is given by16,21,49 _x _y _z\u00f0 \u00deT =R 1 _X _Y _Z T and vx vy vz\u00f0 \u00deT =R 1 vX vY vZ\u00f0 \u00deT \u00f017\u00de MX MY MZ\u00f0 \u00deT =R 1 Mx My Mz\u00f0 \u00deT \u00f018\u00de where R 1 = m1 m2 m3 m4 m5 m6 m7 m8 m9 0 @ 1 A \u00f019\u00de m1 = cos u cos c, m2 = sin c cos u, m3 = sin u, m4 = cos c sin u sin f sin c cos f, m5 = sin f sin u sin c+ cos f cos c, m6 = cos u sin f, m7 = cos c sin u cos f+ sin c sin f, m8 = cos f sin u sin c sin f cos c, m9 = cos f cos u This inverse transformation CTFIB is modeled as shown in Figure 13. It transforms forces from bodyfixed frame to inertial frame and velocities from inertial frame to body-fixed frame. The support beam acts as a base for all the components of the TRMS and all components are mounted on it. The beam is pivoted to the vertical static supporting column as shown in Figure 1. The joint is articulated in such a way that the beam can rotate about the inertial vertical axis (i.e. about the Z-axis or yawing) and the horizontal axis which is normal to the beam axis, that is, the body-fixed y-axis (pitching). Thus, the supporting beam is having 2 degrees of freedom and its configuration at any instant can be defined by yaw angle (c) and pitch angle (u). The rigid body model of the support beam is shown in Figure 17. To get only this 2-DOF motion, four constraints are applied",
            " In the main and tail shield models and counterweight model, rigid joints are modeled with small joint flexibilities to connect those to the support beam. The main and tail rotor BG models are shown in Figures 23 and 24, respectively. The drag and thrust forces due to the rotor are modeled therein as per equations (5)\u2013(9). According to the relative motions between the rotors and the support beam, constraints are imposed at the joints. The main rotor and tail rotors can rotate only about their own axis as shown in Figure 1, with respect to supporting beam. Thus, to allow the rotational motion at the revolute joint of the rotor, pad structure is removed from the model and DC motor models are connected there (see 0 junctions at the bottom left corner of Figure 23 and top right corner of Figure 24). The DC motors apply torque on rotor blades and reaction torques on the support beam. The 0 junction (equal effort and flow sum junction) at the left of the model in Figure 23 indicates that the main rotor angular velocity is the sum of the main motor speed and yaw rate of the support beam (same as shield)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001798_0954406216631370-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001798_0954406216631370-Figure2-1.png",
        "caption": "Figure 2. Double-sided films model in this paper.",
        "texts": [
            " In theory, this concept is also applicable to helical, spiral bevel, and worm gears as long as the geometrical configuration is represented appropriately. In the initial arrangement, the pinion tooth is placed in a central relative position, where the tooth separation is half of the backlash on both side. And the centers of both gears are not allowed to move laterally. Pinion and gear can be modeled as two degree of freedom lumped system with rotational displacements ( g, p) as their coordinates. Figure 2 shows that the proposed double-sided films proposed in this paper, which assumes that the relative angular displacement is equal to zero when the tooth of the secondary gear is just in the center at Purdue University Libraries on June 5, 2016pic.sagepub.comDownloaded from of the gap between the teeth of the primary gear. In such condition, the distance between the flanks of the teeth is indicated with Cb. The lubricant dynamical transmission error x, which represents the motion transfer error along the line of action of gear, can be described as follow x \u00bc Rpp p Rpg g \u00f01\u00de The instantaneous value of lubricant film thickness on the drive side of the distance between the approaching teeth in the normal section is obtained as follows hd \u00f0x\u00de \u00bc Cb x x4 0 Cb x \u00bc 0 Cb \u00fe x x5 0 8< : \u00f02\u00de Then the lubricant film thickness on the coast side is hc \u00bc 2Cb hd",
            " Also, several problems were encountered in this numerical technique. Higher order Runge\u2013Kutta at Purdue University Libraries on June 5, 2016pic.sagepub.comDownloaded from techniques with fixed time step, and even with variable time steps did not improve the result. Part of the problem was attributed to the stiff differential equations, which usually require the use of other types of integration techniques such as Gear\u2019s method for their solution. The computational flow chart is shown in Figure 4. Considering the coordinate systems shown in Figure 2, a positive value of dynamical transmission error corresponds to a positive load between the drive flanks (Fw> 0). Conversely, a negative value of dynamical transmission error with amplitude below the backlash causes a coast flank contact (Fw< 0). In this analysis, the excitation is assumed to be a single harmonic, and the transmission input shaft on which the pinion gears are mounted is subject to nominal speed, superimposed upon engine order vibrations that is only 5% of the nominal speed. To ensure that the numerical simulations and the double-sided film model are reasonable, the first results presented are for the comparison of the shape of the hysteresis loop in Figure 5",
            " Figure 6 also indicates that the amplitude fluctuation of dynamical transmission error could be suppressed with a higher drag torque. Lubricant stiffness is a key parameter in gear dynamics and in determination of the factors such as load-carrying capacity of gear, dynamic tooth loads, and vibration characteristics of the gear systems. In order to quantify this contribution, the effects of the drag torque on the lubricant stiffness must be ascertained. The lubricant stiffness is determined by differentiation the lubricant force with respect to the film thickness.22 According to the position of the oil films in Figure 2, the two springs generated by the driving and coast film act in parallel. Thus, the total gear stiffness of double-sided films is given by K t\u00f0 \u00de \u00bc @Fd w @hd \u00fe @Fc w @hc \u00f08\u00de at Purdue University Libraries on June 5, 2016pic.sagepub.comDownloaded from The lubricant stiffness of single side film and double-sided films alters continually during three mesh cycles with different drag torque is shown in Figures 7 and 8, whose values are far lower than those of solid stiffness in steel\u2013steel contacts, being at least two to three orders magnitudes lower"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000340_978-3-642-27482-4_12-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000340_978-3-642-27482-4_12-Figure1-1.png",
        "caption": "Fig. 1 See through view of the AMiRo robot showing the stackable modules (CAD drawing)",
        "texts": [
            " In Section 4 we describe the software architecture and the required robot programming tools after which we conclude in Section 5 with a brief evaluation and point to further work required. The rules for AMiRESoT robot soccer league require that the soccer robots fit in a vertical cylinder of 110 mm inner diameter [3]. Therefore a cylindrical body is an obvious choice for AMiRo. Cylindrical robot bodies are easy to fabricate from plastic tube material and provide strong structural support with low weight. The top of the cylindrical body can be closed with lids of various shapes. For example a hemispherical lid gives the robot friendly appearance. The inside of this shell (figure 1) has to accommodate the power source (batteries), sensors, actuators (motors and wheels) and the computing hardware. Following the principle of functional modularisation the computing hardware consists of several AMiRo modules (AMs) with a prescribed common electronic interface. Each AM is hosted on its own circuit board and contains its own processing unit that can be a microcontroller, a powerful processor or a programmable device. For best use of space the circuit boards are round and stacked vertically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003374_s00170-020-05807-8-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003374_s00170-020-05807-8-Figure2-1.png",
        "caption": "Fig. 2 a Schematic of experimental set up for micro turning tests, b SLM sample manufactured, and c illustration of cutting parameters",
        "texts": [
            " The hcp pattern can be attributed to both the \u03b1-phase and the\u03b1\u2019martensitic, as they have the same crystalline structure and very similar lattice parameters [15]. No particular heat treatment on Ti6Al4V has been applied after the SLM process. Dry precision turning tests of Ti6Al4V were carried out on a three-axis CNC vertical milling center Roeders RP600 with a spindle speed up to 36,000 rpm. The HSM center is used as a precision turning lathe by positioning a cylindrical titanium sample on the spindle (see Fig. 2a). The sample is composed of two cylinders: a 16 mm diameter allows a clamping by the tool holder and a 20 mm diameter cylinder is composed of a pseudo lattice structure covered by a thin skin of 1.5 mm (see Fig. 2b). The cutting tools fixed on the tool holder were commercial uncoated carbide finishing tools Ifanger Microturn. The K10 micro grains carbide tool is recommended by the manufacturer for precision turning of hard to cut material when precision is needed. An uncoated tool is preferred to limit the cutting-edge radius close to 1 \u03bcm. The tool geometries have been verified on a 3D microscope Alicona Infinite Focus. Four different geometries were used to evaluate the influence of tool geometry on the experimental response (see Table 1). Samples involve dimensional deviations due to process accuracy and roughness due to the thickness of fusion layers and unfused grain. After removing a first thickness of the external surface of the cylindrical parts with a rough tool, dry turning tests were performed with a finishing tool through a length of 4.8 mm with a determined depth of cut. From 1.5 mm, the final skin thickness is then 1 mm (\u00b1 30 \u03bcm) (see Fig. 2c). The cutting process allows a post-processed smooth surface (right) compared to the initial SLM surface (left) (see Fig. 3). The three main cutting forces were measured with a dynamometric table Kistler Minidyn 9256C2 fixed on the machine table and associated with a Kistler amplifier 9017. Data were recorded using a National Instrument CompactDaq 9174 with NI9215 cards and a dedicated Labview program. Fc, Fp, and Ff correspond respectively to force in the table axes X, Y, and Z. The stiffness of the experimental set up leads to a very low force variation during the tests and limited error bars on results",
            " Results were analyzed based on analysis of variance by MINITAB software\u00ae The precision cutting condition leads to minor cutting forces. Due to the high stiffness of the sample and to the low cutting length, no specific vibrations were observed function of the cutting position, contrary to Nieslony et al. [26] when precision turning a 140-mm-long bar of 55NiCrMoV6 bar with an external diameter of 10 mm. A low cantilever allowing a minor force deviation. The maximum value is systematically measured is in the y-axes direction (see Fig. 2c and Table 3), except for test condition n\u00b02. The thrust force can result in a geometrical deviation of the tube during the cutting. However, the higher value is 8.45 \u00b1 0.65 N with the condition n\u00b01, a low force applied on a 20 mm tube reinforced by an internal lattice structure. Concerning the DOE, the main effect of cutting parameters on cutting, feed, and thrust forces are analyzed in Fig. 4. It reveals the particular influence of three parameters. Logically, the depth of cut and feed effects are systematically significant with an increase of force with the increase of values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002959_aeat-05-2019-0094-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002959_aeat-05-2019-0094-Figure1-1.png",
        "caption": "Figure 1 Typical rotor configuration and coordinates",
        "texts": [
            " (2013) developed a general threedimensional dynamic model of helical gear pairs with geometric eccentricity. The gear mesh and bearing flexibility is included in the model as well. Wang et al. (2018) developed an improved time-varying mesh stiffness model of a helical gear pair, in which the total mesh stiffness contains the axial tooth bending stiffness, axial tooth torsional stiffness and axial gear foundation stiffness. This paper used a fixed coordinate system (X,Y, Z) to describe the equation of motion. In the rotor bearing system (Figure 1), U is the axial displacement along the X-axis, V and W indicate the system\u2019s lateral displacement along theY and Z axes. B and C are rotational displacements along the Y and Z, respectively, a is the torsional displacement. The disk in this paper was a rigid body with six degrees of freedom, namely, one axial displacement Ud, two lateral displacements Vd and Wd, and three rotational displacements, ad, Bd and Cd. Because it was a rigid body, the Double-helical geared rotor system Ying-Chung Chen, Tsung-Hsien Yang and Siu-Tong Choi Aircraft Engineering and Aerospace Technology Volume 92 \u00b7Number 4 \u00b7 2020 \u00b7 653\u2013662 strain energy was not taken into account"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003396_med48518.2020.9183165-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003396_med48518.2020.9183165-Figure8-1.png",
        "caption": "Figure 8. The main hardware components of the WSN part: a) the basic sensor node, b) the cluster head.",
        "texts": [
            " The station includes two NTC sensors and two humidity sensors buried in the ground at different depths (50 respectively 30 cm). For sensor nodes we opted to add an 936 Authorized licensed use limited to: University of New South Wales. Downloaded on September 20,2020 at 13:18:15 UTC from IEEE Xplore. Restrictions apply. anemometer and a sensor with multiple functions (temperature, humidity, atmospheric pressure and air quality). The PCB used for developing the UAV payload component is the same as for sensor nodes and cluster head (Fig. 8 and Fig. 9.b). The difference between these three types of WSN elements is the components used for populating each equipment. The payload should communicate with the aerial platform (CAN Messages) and should be able to receive and transmit packets of data to the ground part of the WSN (in this case, the communication is established between payload and CH). To increase the autonomy of stations, the SN and the CH have been equipped with a solar panel which is able to provide up to 5W/h for charging the battery (capacity: 2600mAh and rated voltage: 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000420_math.mag.90.2.87-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000420_math.mag.90.2.87-Figure2-1.png",
        "caption": "Figure 2 The tetrahedron T.",
        "texts": [
            " 2, APRIL 2017 89 According to Theorem 1, to show that a convex body K has the Rupert property, we must exhibit the two projections Po(K) and Pi (K) so that Pi (K) fits in the (relative) interior of Po(K). The tetrahedron, cube, and octahedron can easily be handled geometrically, but the dodecahedron and icosahedron are a bit more difficult. We begin with the tetrahedron. Tetrahedron Let T = ABC D be a regular tetrahedron with unit edge labeled with the equilateral triangle BC D as base and apex A, and take the outer projection Po to be an equilateral triangle C DE of unit side, placed as shown in Figure 2a with E in the plane \u03c0 through the line C D perpendicular to the plane of the base of T. The orthogonal projection of T into \u03c0 nearly fits into the interior of Po, the difficulty being at the vertices C and D. But clearly a small rotation of T about the altitude E M of triangle C E D moves the projections C \u2032 and D\u2032 of C and D on \u03c0 into the edge C D of triangle C E D keeping the projection A\u2032 of A inside Po. It follows that T has the Rupert property, because a small upward translation of the rotated T moves the projections A\u2032, B \u2032, C \u2032, and D\u2032 of all four vertices inside Po. To estimate the Nieuwland constant \u03bd(T) we chose the angle \u03d1 of rotation so that the projected segment A\u2032C \u2032 lies inside Po and parallel to the side EC (Figure 2b). One can show that \u03d1 = 60\u25e6 \u2212 arcsin 1 3 \u221a 6 \u2248 5.264 389\u25e6. (1) Finally, since A\u2032E > C \u2032C = D\u2032 D, C D = 1, and C \u2032 D\u2032 = cos \u03d1 , the least upper bound of the ratio by which T can be expanded and still pass through a similarly situated tunnel in T is This content downloaded from 128.122.230.148 on Thu, 09 Mar 2017 16:20:00 UTC All use subject to http://about.jstor.org/terms C D C \u2032 D\u2032 = sec \u03d1 = 2 5 \u221a 3( \u221a 6 \u2212 1) > 1.004 235, (2) where the surd expression follows from (1). It seems likely that \u03bbT = 2 5 \u221a 3( \u221a 6 \u2212 1), but in any case, \u03bd(T) > 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002579_s11249-016-0801-9-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002579_s11249-016-0801-9-Figure9-1.png",
        "caption": "Fig. 9 Point contact in cylindrical coordinate system (r,h, z) with a the Hertzian contact radius and U(t) the entrainment velocity",
        "texts": [
            " Finally, it provides a complete understanding of the entrapment film mechanisms and gives new insights on the combined effect of deceleration and lubricant rheology. Acknowledgements This work benefits from financial funding of the Agence Nationale de la Recherche via the project Confluence ANR13-JS09-0016-01. The governing equations and assumptions of the problem The Reynolds equation becomes, in cylindrical coordinate system (r,h, z): o oh h3 12g 1 r2 op oh \u00fe 1 r o or r h3 12g op or \u00bc o oh hU\u00f0t\u00de sin h 2r \u00fe 1 r o or hU\u00f0t\u00de r cos h 2 \u00fe oh ot \u00f055\u00de with r the radial position in the contact, a the radius of the Hertzian point contact, as shown in Fig. 9. We assume the problem as axisymmetric despite the existence of a preferred direction\u2014the flow direction\u2014and the existence of the constriction zone. This assumption allows one to simplify the Reynolds equation that becomes: o or r h3 12g op or \u00bc rU\u00f0t\u00de cos h 2 oh or \u00fe r oh ot \u00f056\u00de This equation is first considered in the convergent zone (i.e. for r a and h \u00bc p). Following the same approach and the same assumptions as in Sect. 2.1 and writing h\u00f0r; t\u00de \u00bc h\u00f0a; t\u00de \u00fe hs\u00f0r\u00de, Eq. (56) can be written as, after one integration: oq or \u00bc 6g0U\u00f0t\u00de cos h rhs\u00f0r\u00de R hs\u00f0r\u00de dr r h\u00f0a; t\u00de \u00fe hs\u00f0r\u00de\u00bd 3 \u00fe 6g0 r2 r2 0\u00f0h\u00de r h\u00f0a; t\u00de \u00fe hs\u00f0r\u00de\u00bd 3 dh\u00f0a; t\u00de dt \u00f057\u00de where r0\u00f0h\u00de is the radial coordinate of the position of the maximum pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002129_iccas.2013.6703862-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002129_iccas.2013.6703862-Figure1-1.png",
        "caption": "Fig. 1 The defmition of missile parameters.",
        "texts": [
            " In 978-89-93215-05-295560/13/$15 \u00a9ICROS 50 order to investigate the performance of proposed autopilot, simulation studies with a nonlinear missile model are performed. This paper is constructed as follows: the missile model used in this study is discussed in Section 2. The proposed autopilot is described in Section 3. In Section 4, the performance of proposed method is determined. We conclude this study in Section 5. In this section, the missile model used in this study is discussed. A skid-to-turn (STT) cruciform-type missile model [11], as shown in Fig. 1, is considered. It is assumed that the gravity force can be negligible compared with the commanded acceleration. In addition, the roll can be rapidly stabilized, and then the pitch and yaw motion can be decoupled into two perpendicular channels in STT missile systems. Accordingly, in this study, the missile pitch motion is considered and the equation of motion in the pitch plane can be written as follows: a=- QS [CN (M,a) +CN (M,a) \u00a35]+q mV 0 \" if = \ufffd:d [ Cmo (M, a) + Cmq (M { 2 \ufffd J q + Cm, (M, a) \u00a35 ] (1) aN = :: [ C No (M, a) + C N\" (M, a) \u00a35] where a and q represent the angle-of-attack and the body pitch rate, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001704_978-94-007-6558-0_1-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001704_978-94-007-6558-0_1-Figure4-1.png",
        "caption": "Fig. 4 Possibilities for connecting two simple planetary gear trains in two-carrier compound train. a With two compound and three external shafts. b With two compound and four external shafts. c With one compound and four external shafts",
        "texts": [
            " Ideal external torques are always in a certain ratio, expressed with the torque ratio t TD min : TD max : TR \u00bc T1 : \u00fet T1 : 1\u00fe t\u00f0 \u00deT1 \u00bc 1: \u00fet : 1\u00fe t\u00f0 \u00de \u00f08\u00de in which always TD min\\TD max\\ TRj j \u00f09\u00de however: \u2022 How many degrees of freedom F (F = 1 or F = 2) the gear train is running with; \u2022 Which shaft is fixed at F = 1 degree of freedom; \u2022 What is the direction of power flow, respectively does the gear train works as a reducer or a multiplier at F = 1 degree of freedom, respectively such as a collecting or as a separating gear at F = 2 degrees of freedom, i.e. as a differential; \u2022 Does the gear works as a single or as a component train along with others in a compound planetary gear train. The application of the alternative method is best illustrated with the two-carrier compound planetary gear trains consisting of two of the most commonly used planetary gear trains, shown in Fig. 3. Using the modified symbol of Wolf, the linking up of the two gear trains can be done in different ways, as shown in Fig. 4 with structural schemes\u2014by two or one compound shafts, where we get threeshafts, respectively four-shafts compound planetary gear train. The most commonly used version is the first with two compound shafts and three external shafts (Fig. 4a). As seen from the figure, one of the compound shafts has no outset\u2014it is an internal compound shaft, so that at both ends two equal sized but opposite direction torques are acting. The external torque of the other one, the external compound shaft, which has outset, equals the sum of external torques of the two coupling shafts. The mentioned things above are illustrated by three examples: \u2022 in series connected gear trains (Fig. 5); \u2022 a closed differential gear train with internal division of power (Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000701_tmag.2015.2456338-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000701_tmag.2015.2456338-Figure7-1.png",
        "caption": "Fig. 7. Magnetic flux density of Fig. 4(b) under no current.",
        "texts": [
            " These types are predesigned with the 3D-FEA to generate the maximum suspension force in the same volume size, respectively. Table I shows the specifications of the heteropolar-type MB, the general homopolar-type MB, and the novel homopolar-type MB. From Table I, the magnet volume of the novel homopolar-type MB is decreased to only 7.86% of that of the general homopolar-type MB, because the small magnets are used only for decoupling the suspension forces Fx+, Fx\u2212, Fy+, and Fy\u2212 generated by the four windings, respectively. Fig. 7 shows the magnetic density distribution under no current. The bridges are saturated by inserting the small magnets for blocking the leakage flux. The suspension force and the losses are analyzed. In this paper, the rated suspension current is 3.14 A. In the heteropolar-type and the novel homopolar-type MBs, currents ix+, ix\u2212, i y+, and i y\u2212 are calculated with the following equations: ix+ = ib + ixc, ix\u2212 = ib \u2212 ixc (1) i y+ = ib + i yc, i y\u2212 = ib \u2212 i yc (2) where the bias current ib is for generating the bias flux, and the control currents ixc and i yc are for regulating the suspension force",
            " Regardless of i yc, the suspension forces Fx (i yc = 0 A) and Fx(i yc = 1.57 A) are in good agreement. The maximum decrease ratio of Fx (i yc = 1.57 A) from Fx (i yc = 0 A) is only 6.23% at ixc = 0.314 A. Moreover, with respect to the change of ixc, the variation of Fy(i yc = 1.57 A) is not nearly occurred. Though the width between the magnetic pole is small, the novel homopolar-type is equipped with the superior noninterference characteristic of between Fx and Fy , because of the saturation around bridges, as shown in Fig. 7. In addition, at ixc = i yc = 1.57 A, the maximum suspension force Fb is 8730 N, which is 5.82 times greater than a rotor shaft weight of the envisioned application system. It is known that the novel homopolar-type can generate the sufficient suspension force. Fig. 9 shows the rotor iron loss versus the suspension force Fb plot in these types. The maximum suspension forces of the heteropolar type, the general homopolar-type, and the novel homopolar-type MBs are 8950, 8698, and 8730 N at the same volume size and the same rated current, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002349_978-3-319-45781-9_26-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002349_978-3-319-45781-9_26-Figure1-1.png",
        "caption": "Fig. 1. Mechanical arm structure.",
        "texts": [],
        "surrounding_texts": [
            "The proposed case study is related to the conceptual design of the main body of a flexible robotic arm printed by means of FDM technique. In this paper we do not yet consider the electronic actuation. The arm\u2019s purpose is to support in moving small objects, simulating the shoulder and elbow mechanism. The design requirements are mainly related to the small dimensions of the arm and the coupling parts should give the capability to the arm extremity to reach every point within the volume defined by arm length. Furthermore, the slenderness (1) is considered as the ratio between the sum of the length of all arm components (li) and the average area of arm sections (Am): (1) The robotic arm is divided in two parts, whose movements permit the motion of an object in the 3D space: the mechanical arm and the pliers. The focus of the activity is to determine the optimal design of the structure of both parts, considering manufacturing them as unibody components. The optimal joints play is as- sessed as a balance between the friction necessary to make a position as stable and the needed clearance to allow movements. Robotic mechanism movements are divided into sub-tasks: mechanical arm\u2019s movements, pliers\u2019 movements, pliers\u2019 opening and closing. A kinematical chain has been defined as follows: 3 rotating pairs with parallel lines of actions are used for the mechanical arm movements; 2 rotating pairs with perpendicular lines of action are used for the pliers movements and 2 gearwheels are used for pliers opening and closing. The main selection parameters are precision and ease positioning, component resistance, reduced amount of necessary space, easy architecture. The preliminary design of the mechanical arm and of the plier has been developed, introducing the selected kinematical chains (Figures 1 and Figure 2). The aim of the research activity is to optimize the sizing of the rotating pairs to be manufactured as unibody structures, in order to allow the movements within the joint (clearance effect), with a proper positioning between all the parts and with the feasibility to make all the possible positioning of the arm extremity as stable equilibrium position (posable effect). The final solution is prototyped with a FDM \u2013 Fused Deposition Modeling \u2013 technique as a demonstration of the developed concept."
        ]
    },
    {
        "image_filename": "designv11_30_0002404_mmar.2016.7575286-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002404_mmar.2016.7575286-Figure4-1.png",
        "caption": "Fig. 4. The frame xpyp on the plane ABC",
        "texts": [
            " One of the many input dates of the program dist0 are coordinates of the position of the characteristic points of the object, in the reference frame. On the basis of these coordinates 3 characteristic points the program dist0 calculates the matrix Tob, describing the position and orientation of the frame xobyobzob, associated with the object observed by the cameras. Let\u2019s denote these 3 points by A, B, C. For these points is calculated a matrix Tp, describing the frame xpypzp lying in the plane passing through the points A, B, C. This frame is illustrated in Fig. 4. The origin of the frame xpypzp lies at the point D. This point lies at the intersection of the line AC and perpendicular to this line, passing through the point B. The axis zp is directed upward from point D and is perpendicular to the plane ABC. From the calculations of the program Tobject we known coordinates of the points A, B, C in the frame xyz. The location of these points we will describe by respective vectors Ar  , Br  and Cr  . For calculation of the matrix Tp are needed unit vectors a  , b  and c  of respectively axes xp, yp, zp and the position vector d  of the point D in the frame xyz",
            " Then from (2a, b) were calculated coordinates of the points P1cor\u00f7P4cor in the reference frame xyz. The location of these points in the frame xyz describe the vectors (10). kjir  8372.46344.644735.421 +\u2212\u2212= , kjir  6813.52031.692763.432 ++\u2212= , kjir  7957.50297.684413.403 ++= , kjir  3206.48453.642956.424 +\u2212= . (10) In the program dist0 the points A, B, C are defined as follows: A = P3cor, B = P1cor C =P2cor. For these points the matrix Tp was calculated from the (3a\u00f7d). This matrix describes the frame xpypzp associated with the plane ABC (see Fig. 4, 5). The (11) describes this matrix.             \u2212\u2212 \u2212\u2212 \u2212\u2212\u2212\u2212 = 1000 6850.59999.00063.00014.0 1656.690063.09999.00140.0 5993.400015.00140.09999.0 pT . (11) The matrix pTob describes the object frame xobyobzob from Fig. 7 in the frame xpypzp. From Fig.4 results the frame xpypzp as in Fig. 9, described by (12).             \u2212 = 1000 0100 5.67010 5.42001 ob p T . (12)             \u2212\u2212 \u2212\u2212 \u2212\u2212\u2212 = 1000 3154.59999.00063.00014.0 0780.10063.09999.00140.0 9511.00015.00140.09999.0 obT . (13)             \u2212 = 1000 0100 0001 388010 sT . (14)             \u2212\u2212 = 1000 45010 146100 0001 E . (15) The matrix Tob (see Fig. 5) calculated from (5) has the form (13). For the robot station on which the studies were carried matrix Ts (see Fig.6) has the form (14). The matrix E describing the gripper frame x7y7z7 relative to the working link frame x6y6z6 has the form (15)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000684_978-3-319-15699-6-Figure3.13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000684_978-3-319-15699-6-Figure3.13-1.png",
        "caption": "Fig. 3.13 Formations used in the experiment. a Pyramidal. b Ring. c Star Ring. d Complete",
        "texts": [
            " 44 3 The Data Association Problem We have also tested our proposal with real images considering different scenarios such as teams of mobile robots with cameras or surveillance camera networks. In each example we have used different features and functions to find the local correspondences. In the first experiment there are 6 robots moving in formation (around 5m away from each other). Each robot acquires one image with its camera and extracts SURF features [11] (Fig. 3.12). The epipolar constraint plus RANSAC [48] is used for the local matching. The detection and resolution of inconsistencies is analyzed for four different typical communication graphs (Fig. 3.13). The error function used for the Maximum Error Cut algorithm is the Sampson distance. We have chosen man made scenarios to be able to manually classify the matches, see Fig. 3.12. Although ground truth is not available in this example, by looking at the correspondences we have counted the amount of full matches. This number is very small or even zero due to missing matches and occlusions caused by the trees 3.5 Experiments 45 in the images. For that reason we also define a partial match when 3 or more robots correctly match the same feature, because in such case the propagation is required for the association"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002234_053001-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002234_053001-Figure1-1.png",
        "caption": "Figure 1. Three basic ways forces can be applied to deform a solid. (a) Tension; (b) compression; (c) shear. F: force; F1, F2: shearing forces (see text).",
        "texts": [
            " Most biological solid materials, however, have at least some degree of fluid-like behavior and are thus technically viscoelastic; because viscoelastic properties combine solid and fluid aspects, we need to understand both conventional solids and conventional fluids before tackling viscoelastic solids in section 4. Nevertheless, the hardest biological solids do act much like engineering solids, and the sophisticated methods that engineers have developed to analyze solid materials can be usefully applied to many stiff biological materials. Solids can be deformed in different ways. A rope supporting a weight is loaded in tension, where the ends are being pulled apart. The concrete floor I stand on is loaded in compression, where the top and bottom surfaces are being pushed closer together (figure 1). Shearing is when forces are applied in such a way that they tend to make one part of the object slide past another part. Consider the rectangular solid in figure 1(c): force F1 pushes on the upper left while force F2 pushes on the lower right, so that the upper part of the object tends to slide to the right and the lower part tends to slide to the left. Some materials are flexible, and can resist tension but not compression, hence the old saying, \u2018you cannot push on a rope\u2019. Others resist compression mightily\u2014they are \u2018hard\u2019 in the conventional sense\u2014but do not resist tension well, such as concrete. Many biological materials are somewhere in between, with a moderate amount of tensile and compressive strength and some flexibility; Wainwright et al (1982) call these materials \u2018pliant\u2019, to distinguish them from materials they categorize as \u2018tensile\u2019 and \u2018rigid\u2019",
            " The most obvious source of drag is viscosity; after all, we define a fluid as a material that resists rate of shear, and viscosity is a measure of this resistance. Viscosity is given by t d tan d , 16t m g = ( ) ( ) where \u03c4 is the shear stress and tan(\u03b3) is the shear strain (as in (4)) and \u03bc is the coefficient of viscosity, also called the dynamic viscosity. Unless otherwise specified, \u2018viscosity\u2019 is usually taken to mean \u2018dynamic viscosity\u2019, and I will follow that convention. A useful simplification uses the same shape shown being sheared in figure 1(c). Imagine this shape represents a region of fluid. If the surface area of the top and bottom are equal, the distance between top and bottom is d, and the top moves past the fixed bottom at speed v, then the shear stress can be written: v d . 17t m = ( ) Equation (17) and arrangements of plates or cylinders sliding past each other are the basis for common devices used to measure viscosity. Dynamic viscosity is a material property. If the viscosity of a fluid does not change with shear rate, the fluid is called \u2018Newtonian\u2019; air and water are both Newtonian fluids"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000658_1.4030937-Figure9-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000658_1.4030937-Figure9-1.png",
        "caption": "Fig. 9 Three-dimensional models created by SOLIDWORKS",
        "texts": [
            " According to the given mean point of the generating gear and the reference point of the cutter, the virtual machine settings for cutting right and left flanks of tooth surfaces, as applied to this case, are as listed in Table 2. Substituting the cutter parameters and the virtual machine settings into Eqs. (8)\u2013(10) yields the tooth surface positions and normals of the pinion and gear listed in Table 3. Once solved, the topographical points are used to build the 3D models (created in SOLIDWORKS) shown in Fig. 9. The correlation between the generated lengthwise crowning and the blade profile angle is shown in Fig. 10: The amount of crowning is almost proportional to the blade profile angle. The contact performance of the numerical gear pair is further assessed using ease-off and TCA, whose results are shown in Figs. 11 and 12, respectively. Both these analytic tools were specially developed for the purpose in our research laboratory. According to the ease-off results, lengthwise crowning is produced by the cutter\u2019s profile angle and the convex tooth surfaces generated are advantageous in absorbing assembly errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure4.19-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure4.19-1.png",
        "caption": "FIGURE 4.19 Capacitor with an electrode rotated about its center.",
        "texts": [
            " This is approximately a 10% drop\u2014this may or may not be important, depending on the application. The final variation to be presented in this section is the calculation for a rotated electrode. The calculation is not difficult or involved; it was omitted in the original program because it might have been a distraction. The following equations rotate a point (xj,yj,zj) about the z axis by an angle \u03b8: xnew = xj cos \u03b8\u00f0 \u00de\u2212yj sin \u03b8\u00f0 \u00de \u00f04:7\u00de ynew = xj sin \u03b8\u00f0 \u00de+ yj cos \u03b8\u00f0 \u00de \u00f04:8\u00de This code is implemented in get_data.m. It is not executed until a nonzero rotation is encountered Figure 4.19 shows an example of an electrode rotating about its center. The data file used is shown in Table 4.4. Figure 4.20 shows the result of this rotation. 84 Examples Using the Method of Moments electrode), it is necessary only to define that point as the center of the coordinate system. TABLE 4.4 Data File for Capacitor Shown in Figure 4.19 0.01 0 0 25 200 0 \u22120.5 45 0 0 25 200 0.1 0.5 0 260 240 220 200 180 160C ( pf d) 140 120 100 80 0 20 40 60 80 \u03b8 (\u00b0) 100 120 140 160 180 FIGURE 4.20 Capacitance as a function of rotation angle \u03b8 for capacitor shown in Figure 4.19. 4.8 Varying the Geometry 85 Figure 4.21 shows an example of this situation; with respect to the origin, the lower left corner of the electrodes is (\u22120.25,\u20130.25). The data file is shown in Table 4.5, and the capacitance as a function of \u03b8 is shown in Figure 4.22. Figure 4.22 shows a curious capacitance behavior\u2014while it is expected that the capacitance will fall drastically as \u03b8 is increased from 0, the capacitance variation TABLE 4.5 Data File for Capacitor Shown in Figure 4.22 0.01 \u22120.25 \u22120.25 20 160 0 \u22120"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002254_s0081543816060171-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002254_s0081543816060171-Figure8-1.png",
        "caption": "Fig. 8. Realization of the system describing the dynamics of a Chaplygin ball with Suslov\u2019s constraint.",
        "texts": [
            " For the parameters corresponding to Fig. 7b, we have two limit cycles: a stable cycle and an unstable one. Then, as the parameter a decreases, the cycles disappear. 3. A CHAPLYGIN BALL WITH SUSLOV\u2019S CONSTRAINT The paper [11] is concerned with a system that is equivalent to the problem of the motion of a Chaplygin ball with the additional Suslov constraint (2.1). Moreover, this paper proposes an implementation of this system that allows one to construct another possible nonholonomic generalization of the Euler\u2013Poisson equations (see Fig. 8). In this case, as in Vagner\u2019s implementation [50], it is assumed that the rigid body B is equipped with wheels (on one axis) and enclosed in a fixed spherical shell. The condition that there be no slipping in the direction perpendicular to the plane of the wheels leads to the Suslov constraint (\u03c9, e) = 0, where \u03c9 is the angular velocity of the body and e is the body-fixed vector lying in the plane of the wheels perpendicularly to the axle supporting the wheels. Below we make use of a body-fixed PROCEEDINGS OF THE STEKLOV INSTITUTE OF MATHEMATICS Vol",
            " At the point P the cavity is in contact with a freely rotating homogeneous ball S whose center is fixed. At the contact point P , the no-slip condition (mutual spinning is not prohibited) is satisfied: R\u03c9 \u00d7 \u03b3 = RS\u03c9S \u00d7 \u03b3, where \u03c9S is the angular velocity of the ball and \u03b3 is the unit vector directed along the axis joining the centers of the cavity and the ball. If in this implementation we choose the fixed ball S inside the cavity in such a way that the vector joining its center with the center of the body\u2019s cavity is vertical (see Fig. 8), then the equations of motion of the body B in the body-fixed coordinate system take the form I\u0302 ( \u03c9\u03071 \u03c9\u03072 ) = ( \u2212(I13\u03c91 + I23\u03c92)\u03c92 + b3\u03b32 \u2212 b2\u03b33 (I13\u03c91 + I23\u03c92)\u03c91 + b1\u03b33 \u2212 b3\u03b31 ) , \u03c93 = 0, \u03b3\u03071 = \u2212\u03b33\u03c92, \u03b3\u03072 = \u03b33\u03c91, \u03b3\u03073 = \u03b31\u03c92 \u2212 \u03b32\u03c91, I\u0302 = ( I11 +D(\u03b322 + \u03b323) \u2212D\u03b31\u03b32 \u2212D\u03b31\u03b32 I22 +D(\u03b321 + \u03b323) ) , D = R2 R2 S IS , (3.1) where IS is the moment of inertia of the ball S, Iij are the components of the body\u2019s tensor of inertia relative to the point O, and the axes of the coordinate system have been chosen in such a way that I12 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000008_rusautocon.2019.8867636-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000008_rusautocon.2019.8867636-Figure2-1.png",
        "caption": "Fig. 2. DJI F450 Quadrotor Frame",
        "texts": [
            " PARAMETERS IDENIFICATION In this section we discuss the identification of some parameters of the model given by the previous equations. If we neglect the dynamics of the rotors, these parameters can be divided into two types: static (aerodynamic constant of lift b and drag d) and dynamic (mass and moments of inertia). The constants of aerodynamic friction (Kfax, Kfay, Kfaz) will be neglected in what follows, since they are not necessary for diagnosis and the synthesis of the control laws. A modified DJI F450 quadrotor frame is used in our experiment (see Fig. 2). In this experiment the quadrotor was fixed to a mass whose value was chosen so as to prevent the quadrotor from taking off when the rotors are rotating at full speed. The scale was reset, then we varied the speed of each rotor (increasing the duty cycle of the control signal) and we noted the value of the corresponding mass (see Fig. 3); from which we have calculated the value of the lift force: \ud835\udc39\ud835\udc39 = 4\ud835\udc4f\ud835\udc4f\ud835\udc64\ud835\udc642 = \ud835\udc43\ud835\udc43 = \ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a The slope of this curve is equal to 4.4154.10-5 N. (rad / s)-2, which represents the lift constant multiplied by 4; the lift constant is: b=1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000511_j.orgel.2015.08.027-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000511_j.orgel.2015.08.027-Figure2-1.png",
        "caption": "Fig. 2. The illustration shows the setup of the electro-optical measurement system. The L directions for the operation of a normally white TN mode.",
        "texts": [
            "2 mm thickness indium tin oxide (ITO) coated glass sandwich cells with mutually perpendicular rubbed alignment layers. To observe the two kinds of structure of the gels in the cells with anti-parallel rubbed alignment layers and mutually perpendicular alignment layers, the dark-field mode of Olympus BX51 optical microscope (OM) was conducted. The detailed electro-optical properties of the supramolecular LC gels with different concentration of F8BT in TN cell were further measured by our experiment system shown in Fig. 2. The system is composed of a HeeNe laser, a pair of polarizers, sample cell mounted in rotatable holder, and the photodetector. An unpolarized HeeNe (632.8 nm) laser was used as an incident light source. At the beginning, the polarizers were tuned to be perpendicular to each other to make sure the transmittance minimum. Then the sample cell with LC gels was placed between the crossed polarizers and tuned to make the transmittance reach maximum. And the sample cell was driven by an AC field (1 kHz, squarewave) supplied by the function generator SRS DS360 which was controlled by a computer with LabView programming",
            " This high PL ratio finding shows that the main-chain direction of F8BT molecules should align with the rubbing direction of cell, that is, the long-axis direction of LC molecules, which is consistent with results of the OM picture and the phase diagram since the Tiso-ne is higher than Tsol-gel. And the polarized PL spectra for samples prepared in cells with mutually perpendicular alignment layers are shown in Fig. 5(b). It can be seen that the polarized PL spectra in two directions are almost the same. This result unambiguously indicates that the F8BT molecules would not have preferred orientation in the TN cells. And then the electro-optical properties of the supramolecular LC gels in TN cells were examined by the HeeNe Laser experimental system shown in Fig. 2. The sample cells filled with the gels in different concentrations of the gelator F8BT were prepared through the previously mentioned experimental procedures. Fig. 6 shows the voltage-transmittance (V-T) curves of the LC gels with different concentrations (0, 0.1, 0.2, and 0.4 wt%) of the gelator F8BT in TN cells. As the electric field is off, the gels present the bright transmission state. When the electric field turns on, the transmittance starts to decrease. From the V-T curves, a bump peak found clearly in transmittance in the case of pure LC 5CB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001040_j.mporth.2016.05.014-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001040_j.mporth.2016.05.014-Figure7-1.png",
        "caption": "Figure 7 (a) Rolling contact occurs when there is no relative motion between the two surfaces. The point where the wheel touches the ground is the instantaneous centre of rotation since there is no motion at this point. (b) Sliding contact occurs when there is no relative translation of the two objects. The instantaneous centre of rotation is therefore located in the centre of the wheel.",
        "texts": [
            " In three dimensions, there are six possible degrees of freedom: three translational along three perpendicular axes and three rotational about each of the three axes (Fig. 6). Relative motion at the joint surfaces In the diarthrodial joints, the relative motions of the bones are constrained by the geometry of the joint surfaces and action of the ligaments and muscles spanning the joint. When the two 2005 Elsevier Ltd. All rights reserved. joint surfaces remain in contact, they may move relative to each other by rolling or sliding. Figure 7 shows the simple case of a circular wheel on a flat surface. Rolling occurs when there is no relative velocity, that is, no slip, between the two contacting points and the ICR is located at the point of contact. Sliding contact occurs when there is no resistive force between the two surfaces and the ICR is located at the centre of the wheel. The relative motion across the human joints is generally a combination of rolling and sliding. Both take place simultaneously in the knee joint, whereas in the hip and shoulder joints, sliding motion predominates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002047_s12206-015-1149-z-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002047_s12206-015-1149-z-Figure6-1.png",
        "caption": "Fig. 6. Temperature distributions: (a) in a solid structure; (b) of the cooling flows in the high-speed PM motor at a rotor speed of 35 krpm with no cooling air.",
        "texts": [
            " Note that the rotor is \u201cthermally floating\u201d without strong heat sinks. In practice, maintaining the rotor temperature steady and low is very important from the viewpoint of rotor dynamics and lubrication in high-speed rotating machines. In worst cases, excessive thermal expansions caused by elevated rotor temperature can cause rotor-bearing system failure due to thermal seizures. Fig. 5 shows the temperature distribution (a) in a solid structure and (b) of the cooling flows in a high-speed PM motor at a rotor speed of 35 krpm for a sine-wave current input. Fig. 6 shows the temperature distribution (a) in a solid structure and (b) of the cooling flows in a high-speed PM motor at a rotor speed of 35 krpm without cooling air. Note that only small changes in the flow pressure and velocity of the cooling flows are observed; hence, they are omitted for simplicity. A further study is conducted with the reference model using the measured input current on a PM motor that drives a turbo compressor at a rotor speed of 60 krpm (see Appendix for the measured current and loss analysis)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001380_icra.2014.6907727-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001380_icra.2014.6907727-Figure1-1.png",
        "caption": "Fig. 1. Schematic used for kinematic model of a 3-DOF planar cannula",
        "texts": [
            " There are two slots along the length of each straight segment and two antagonistic SMA actuators are placed at each joint. The SMA actuators are 0.53 mm in diameter and they are annealed in an arc shape. In this work, all SMA actuators lie in a plane. Hence, the cannula moves in a plane perpendicular to the location of the SMA actuators. The joints are covered with non-conductive rubber sheath for heat isolation and electrical insulation of the SMA actuators inside soft-tissue. The schematic of the cannula is shown in Fig. 1. There are 3 degrees-of-freedom (DOF): One DOF at each joint and one insertion DOF. The configuration of the cannula in a plane can be described using Eq. 1. x = L1 + u+ L2cos\u03b11 + L3cos\u03b2 + r1sin\u03b11 + r2(sin\u03b12cos\u03b11 \u2212 sin\u03b11(1\u2212 cos\u03b12)) y = L2sin\u03b11 + L3sin\u03b2 + r1(1\u2212 cos\u03b11) + r2(sin\u03b11sin\u03b12 + cos\u03b11(1\u2212 cos\u03b12)) \u03b2 = \u03b11 + \u03b12 (1) The joint angles, \u03b11 and \u03b12, are shown in Fig. 1. L1, L2, L3 are the lengths of the straight segments from base to tip. The position of the tip is given by (x, y) and \u03b2 is the orientation of the tip. The radius of curvature of the joints are given by r1 and r2. The relation between the arc radius, r, and joint angle, \u03b1, is r = \u2113/\u03b1 where \u2113 is the length of the SMA actuator between consecutive links. The three equations representing the geometry of the cannula is solved using a 4th order Runge-Kutta solver. The algorithm was implemented in C++. The experimental setup used for ultrasound guidance is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001447_cjme.2014.1118.170-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001447_cjme.2014.1118.170-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram and coordinate for turbulent journal bearing",
        "texts": [
            " The Reynolds equation of a finite length turbulent journal bearing with couple stress is solved based on Sommerfeld number and Ocvirk number combing with the multi-parameter principle. The influences of eccentricity ratio and width-to-diameter ratio on oil film load-carrying capacity are analyzed. The proposed method can provide reference for the design of the rotor system supported by a turbulent journal bearing with couple stress. 2 Reynolds Equation under Turbulent Condition with Couple Stress Fluid Fig. 1 shows the schematic diagram and the coordinate of the turbulent journal bearing. Fig. 2 shows the cross section of the turbulent journal bearing. In Fig. 1, Ob is the bearing center, Oj is the journal center,  is the angle determined from negative y direction to the oil film location in the clockwise direction, is the deviation angle,  is the angle determined from OjOb to the oil film location in the clockwise direction. fr and ft are nonlinear oil film forces acting on journal in the radial and tangential directions respectively. fx and fy are nonlinear oil film forces acting on journal in the x and y directions respectively. h is the oil film thickness, \u03c9 is the angular velocity of the rotor, R is the radius of the bearing, r is the journal radius, and mg is the bearing load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003116_calcon49167.2020.9106511-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003116_calcon49167.2020.9106511-Figure1-1.png",
        "caption": "Fig. 1. Typical deep-groove ball bearing [5]",
        "texts": [
            " Parallelly applied force to the shaft is named axial force and sheer applied force to the shaft is understood as radial force. So, for up the lifetime of bearings the foremost small print to appear once of this instrumentation are correct alignment, correct placement and enough lubrication [7]. ISBN: 978-1-7281-4283-8 502 PART NO.: CFP20O01-ART Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 05,2020 at 01:16:31 UTC from IEEE Xplore. Restrictions apply. An example of a typical deep groove bearing is given in Fig. 1. In Fig. 2. we can see that a needle bearing consisting of an outer race, an inner race, balls, a shaft and a cage holding the balls. In the figure, it is often ascertained that the zone of load and the distribution of load are given along the direction of applied force. Virtually, the outer race is the stationary part and the inner race and the balls are the rotating part. As the outer race is exposed to load, cracks and pits may happen. The rotating parts may contain other defects due to constant rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure22-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure22-1.png",
        "caption": "Figure 22. Speed field contours for 700 rpm.",
        "texts": [],
        "surrounding_texts": [
            "Table 2 and the graph (Figure 19) show the calculations of the maximum speed of the velocity fields for a fixed drum diameter of 650 mm and different values of the drum rotation speeds. The diameter and rotation speed of the beater are equal to the initial values. [22] ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0002654_cipech.2016.7918731-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002654_cipech.2016.7918731-Figure2-1.png",
        "caption": "Fig. 2. The quad-rotor angle movement",
        "texts": [],
        "surrounding_texts": [
            "Genetic Algorithm Based Trajectory Stabilization of Quadrotor\nMukul Gaur a, Himanshu Chaudhary b, Shahida khatoon c, Ravindra Singhd\naDCRUST, Murthal, India b, cDepartment of Electrical Engineering, Jamia Millia Islamia, New Delhi-25, India\ndScientist \u2018G\u2019, LASTEC, DRDO, New Delhi, India himanshuchaudhary.h@gmail.com\nAbstract - This paper deals with control and stabilization of a quadcopter UAV using Genetic algorithm tuned PID controller. In this work, conventional GA is improved in two ways: firstly, crossover fraction and mutation rate is made adaptive by using a Fuzzy logic controller. Secondly, an advanced randomness is provided in GA by changing half of its initial population with random candidates after a fixed generations. Simulation results proven to be more optimized with the proposed controller in respect to the both transient response and robustness in presence of adverse condition or disturbances.\nKeywords \u2014 Quadcopter, Dynamic Modeling, PID Controller, Genetic Algorithm (GA). Modified Genetic Algorithm.\nI. INTRODUCTION Unmanned aerial vehicles (UAVs) are capable of flight with no pilot on-board. They can be controlled remotely by an operator or autonomously via pre-programmed flight paths. These are used by the military, as tools for search and rescue operations, warrant continued development of UAV technology. Recently, Non-military use of UAVs has also increased. Some of them are: Pipeline/power line inspection, border patrol, search and rescue, oil/natural gas searches, fire prevention, wild life management, aerial photography and film making, topography and agriculture etc. [1]\nLift of quadcopter is generated by four motors. Control of such a craft is accomplished by varying the speeds of these motors relative to each other. Hence, it demands a sophisticated control system to allow for a balanced flight, as the dynamics of such a system require constant adjustment of four motors simultaneously. The goal of this work is to design a robust and efficient control system for quad-rotor controlling.\nFor this purpose, a number of control strategies have already been used. The main quadrotor control methodsused in the literature are:Linear Quadratic Regulator (LQR) [2], proportional \u2013 integral \u2013 differential (PID) [2], control methods based on adaptive techniques, backstepping control [3], dynamic feedback [4],control strategy based on visual feedback [5] etc.. Some of these require lot of computation while other need an exact model of the quadcopter. In this paper, we have used PID controller for quadcopter control. PID controller has simple structure &its performance has already been tested for various processes. Its parameters also can be easily tuned for optimal performances [6].\nGenetic algorithm GA is a versatile method in the field of optimization over the past few years& have already been used\nto solve the optimization problems with minimal problem information[7] e.g. Task of tuning of PID controller [8], evolving the structure of Artificial Neural Networks [9] & evolving the structure of Fuzzy Logic Controllers [10], Economic Load Dispatch Problems [11] & many other optimization problems. Elitism, arithmetic crossover & mutation operators are main operations of GA.\nGAs perform global searches, but their long computation times limit them when solving optimization problems on large scale. In this work, an improved Genetic Algorithm has been proposed. Fuzzy logic controllers have already been used to get adaptive crossover and mutation rates of GA [12].The same technique is being used in this work. But, the structure of the FLCs used is different w.r.t the membership functions. Moreover, to enhance the randomness in the algorithm, half the initial population is generated randomly after five generations [9]. Results show an up gradation in over-all performance using the designed algorithm. The paper is organised as follows: Section I deals with the Introduction and diverse application of Quad-rotor, In section II, Quad-rotor model along with its complex dynamics is presented, Section III presents modified genetic algorithm integration with designed Quad-rotor model, followed by simulation results study & concluding remarks at last.\nII. QUADCOPTER MODEL In this section, a brief summary of quadrotor model used in this work has been discussed. Figure 1 shows the theoretical model of quadcopter. Quadcopter is multi-input multi-output dynamical nonlinear system. A highly complex set of equations with their derivations can be found in [1]. Also simplification of these equations is presented in [13]. Each rotor produces moments and vertical force. These moments have been observed to be linearly dependent on the forces at low speeds. Since there are four input forces and six output states (x, y, z, \u03c8, \u03b8, \u03c6) the quad-rotor is an under actuated system and thus modelling a vehicle such as quadrotor is not an easy task. Therefore, the aim is to develop the model as much realistic as possible. The sum of the thrust of each rotor is the collective input (U1). As, the input is the thrust generated by the four rotors which are fixed, different movements of the quadcopter are obtained by increasing or decreasing the speeds of various motors. The quad-rotor dynamics are determined from a set of equations of motion. The complexity of equations increases with increase in accuracy. There are six degrees of freedom of a rigid body movement: of which three define the reference position\nGenetic Algorithm Based Trajectory Stabilization of Quadrotor\nMukul Gaur a, Himanshu Chaudhary b, Shahida khatoon c, Ravindra Singhd\naDCRUST, Murthal, India b, cDepartment of Electrical Engineering, Jamia Millia Islamia, New Delhi-25, India\ndScientist \u2018G\u2019, LASTEC, DRDO, New Delhi, India himanshuchaudhary.h@gmail.com\nAbstract - This paper deals with control and stabilization of a quadcopter UAV using Genetic algorithm tuned PID controller. In this work, conventional GA is improved in two ways: firstly, crossover fraction and mutation rate is made adaptive by using a Fuzzy logic controller. Secondly, an advanced randomness is provided in GA by changing half of its initial population with random candidates after a fixed generations. Simulation results proven to be more optimized with the proposed controller in respect to the both transient response and robustness in presence of adverse condition or disturbances.\nKeywords \u2014 Quadcopter, Dynamic Modeling, PID Controller, Genetic Algorithm (GA). Modified Genetic Algorithm.\nI. INTRODUCTION Unmanned aerial v hicl (UAVs) are capable of flight with no pil t on-boa d. They can be controlled remotely by an operator or autonomously via pre-programmed flight paths. These are used by the military, as tools for search and rescue operations, warrant continued development of UAV technology. Recently, Non-military use of UAVs has also increased. Some of them are: Pipeline/power line inspection, border patrol, search and rescue, oil/natural gas searches, fire prevention, wild life manageme t, aerial photography and film making, t pography and agriculture etc. [1]\nLift of quadcopter is generated by four motors. Control of such a craft is accomplished by varying the speeds of these motors relative to each other. Hence, it demands a sophisticated control system to allow for a balanced flight, as the dynamics of such a system require constant adjustment of four motors simultaneously. The goal of this work is to design a robust and efficient control system for quad-rotor controlling.\nFor this purpose, a number of control strategies have already been used. The main quadrotor control methodsused in the literature are:Linear Quadratic Regulator (LQR) [2], proportional \u2013 integral \u2013 differential (PID) [2], control methods based on adaptive techniques, backstepping control [3], dynamic feedback [4],control strategy based on visual feedback [5] etc.. Some of these require lot of computation while ot er need an exact odel of the quadcopter. In this paper, w hav used PID controller for quadcopter control. PID controller has simple structure &its pe formance has already been tested for various processes. Its parameters also can be easily tuned for optimal performances [6].\nGenetic algorithm GA is a versatile method in the field of optimization over the past few years& have already been used\nto solve the optimization problems with minimal problem information[7] e.g. Task of tuning of PID controller [8], evolving the structure of Artificial Neural Networks [9] & evolving the structure of Fuzzy Logic Controllers [10], Economic Load Dispatch Problems [11] & many other optimization problems. Elitism, arithmetic crossover & mutation operators are main operations of GA.\nGAs perform global searches, but their long computation times limit them when solving optimization problems on large scale. In this work, an improved Genetic Algorithm has been proposed. Fuzzy logic controllers have already been used to get adaptive crossover and mutation rates of GA [12].The same technique is being used in this work. But, the structure of the FLCs used is different w.r.t the membership functions. Moreover, to enhance th randomness in the alg rithm, half the initial population is enerated randomly after five gen rations [9]. Results show an up gradation in over-all perf rmance using the designed lgorithm. Th paper is organised as follows: Section I deals with the Introduction and divers application of Quad-rotor, In section II, Q ad- ot r model along with its complex dynamics is presented, Section III presents modifi d genetic algorithm integration with d sig ed Quad-rotor model, followed by si ulation results study & concluding remarks at last.\nII. QUADCOPTER MODEL In this section, a brief summary of quadrotor model used in this work has been discussed. Figure 1 shows the theoretical model of quadcopter. Quadcopter is multi-input multi-output dynamical nonlinear system. A highly complex set of equations with their derivations can be found in [1]. Also simplification of these equations is presented in [13]. Each rotor produces moments and vertical force. These moments have been observed to be linearly dependent on the forces at low speeds. Since there are four input forces and six output states (x, y, z, \u03c8, \u03b8, \u03c6) the quad-rotor is an under actuated system and thus modelling a vehicle such as quadrotor is not an easy task. Therefore, the aim is to develop the model as much realistic as possible. The sum of the thrust of each rotor is the collective input (U1). As, the input is the thrust generated by the four rotors which are fixed, different movements of the quadcopter are obtained by increasing or decreasing the speeds of various motors. The quad-rotor dynamics are determined from a set of equations of motion. The complexity of equations increases with increase in accuracy. There are six degrees of freedom of a rigid body movement: of which three define the reference position\nMukul Gaur DCRUST, Murthal, India\nHimanshu Chaudhary Department of Electrical Engineering,\nJamia Millia Islamia, New Delhi\u201325, India himanshuchaudhary.h@gmail.com\nShahida Khatoon Department of Electrical Engineering,\nJamia Millia Islamia, New Delhi\u201325, India\nNew Delhi, India\n978-1-4673-9080-4/16/$31.00 \u00a92016 IEEE",
            "(usually the centre of mass), and the other three defines the orientation of the vehicle.\nFig.1. Theoretical Model of Quadcopter\nwhere, X= (x, y, z) is the position vector of the quad-rotor. v= (u, v, w) is the speed vector of the quad-rotor. \u03ac= (\u03c8,\u03b8,\u03c6) the Euler angles (for pitch, roll and yaw respectively). \u03c9 = (p, q, r) s the angular speed vector. Two coordinate systems are required to define at any time the instantaneous state of the platform. First, a fixed system with the x-axisalong the front, the z-axis down and the y-axis to the right of the craft. Second, an earth fixed inertial system using the convention type of aviation applications. The one frame rotation relative to the other can be described by rotation matrix, which comprise of three independent matrices and describes the rotation of craft about every single earth frame axes. These rotation matrices can be written as follows.\n\ufffd\u2205 = \ufffd 1 0 0 0 \ufffd\ufffd\ufffd\u2205 \ufffd\ufffd\ufffd\u2205 0 \ufffd\ufffd\ufffd\ufffd\u2205 \ufffd\ufffd\ufffd\u2205 \ufffd\n\ufffd\ufffd = \ufffd \ufffd\ufffd\ufffd\ufffd 0 \ufffd\ufffd\ufffd\ufffd\ufffd 0 1 0\n\ufffd\ufffd\ufffd\ufffd 0 \ufffd\ufffd\ufffd\ufffd \ufffd (1)\n\ufffd\ufffd = \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd 0 \ufffd\ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd 0 0 0 1 \ufffd\nThe total rotation matrix equation can be calculated by product of above matrices:-\n\ufffd = \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd (2)\nThis gives the resultant as shown in the equation 3.\n\ufffd\ufffd\ufffd\ufffd = \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd \ufffd (3)\nWhere S\u03b8 = Sin (\u03b8), C\u03c8 = Cos (\u03c8), and R is the matrix transformation. Using equation (3), we can rotate the quadcopter around x, y & z axis. In this work, a simplified model of quad-copter is used. While modelling the quad-copter following assumptions are being made:\n1. The structure of quad-rotor is symmetrical and rigid. 2. The value of Inertia matrix (I) of vehicle is negligible and to be neglected. 3. The propellers are not flexible. 4. Drag and thrust are proportional to the square of the propellers speed. The equation of motion can be written using the force and moment balance as represented by equation 4.\nx\ufffd = (u\u2081(cos sin cos \ufffd sin sin) \ufffd \ufffd\u2081x\ufffd\n\ufffd\ufffd = (\ufffd\u2081(sin  \ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd \ufffd\ufffd\ufffd) \ufffd \ufffd\u2082\ufffd\ufffd )/\ufffd (4)\nz\ufffd = (u\u2081(cos cos\ufffd \ufffd) \ufffd \ufffd\u2083z\ufffd)/m\nThe Ki's are the drag coefficients. Here, drag is assumed to be zero as drag is negligible at low speeds. When the centre of gravity moves, the angular acceleration becomes less sensitive to the forces, therefore stability is increased. This will decrement the roll and pitch moments and also the total vertical thrust. The inputs can be represented by equation. 5.\nU1= (Th1 + Th2 + Th3 + Th4)/m\nU2= l (-Th1 - Th2 + Th3 + Th4)/I1\nU3= l (-Th1 + Th2 + Th3 - Th4)/I2 U4= C (Th1 + Th2 + Th3 + Th4)/I3\nEuler angles equations are given by equation. 6\n\ufffd = u\u2082 \ufffd \ufffd\ufffd\ufffd\ufffd /I\u2081 \ufffd = u\u2083 \ufffd \ufffd\ufffd\ufffd\ufffd /I\u2082 \ufffd = u\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd /I\u2083 Here, Thi, for i = 1 to 4 are the thrust of four motors of quadcopter. Ii, for i = 1 to 3 are the moments of inertia of the craft with respect to the axes. Ki, for i= 1 to 3 are the drag coefficients with centre of gravity assumed to be on the origin. All states cannot be controlledsimultaneously. A possible combination of controlled outputs can be x, y, z and a good controller must reach a desired yaw angle and position while keeping the pitch and roll angles constant. Theta (\ufffd\ufffd) and Pitch (\ufffd\ufffd) can be extracted by the equations 7 & 8 respectively.\n(5)\n(6)",
            "\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd (7)\n\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd\ufffd \ufffd \ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd\ufffd\ufffd\ufffd\ufffd\ufffd\ufffd \ufffd (8)\nThe overall Simulink model of the quadcopter model is shown in Fig. 3.\nIII. MODIFIED GENETIC ALGORITHM\nGAs are based on natural genetics and selection. These are random but intelligent search techniques used to solve optimization problems.Basic flow chart of GA consists of following steps :- 1. Population is randomly initialized 2. Fitness of population is determined. 3. Above two steps are repeated again and again. 4. Select parents from population. 5. Perform crossover & mutation on parents creating next\npopulation 6. Step 2 to 5 are repeated until best individual is good\nenough (satisfying termination criteria). Further ability of canonical genetic algorithm can be improved by modifying it as follows: 1. Adaptation of crossover and mutation rate. 2. Enhancement of element of randomness. To increase the speed of convergence of GA, crossover fraction and mutation rates are made adaptive and for this, two fuzzy logic controllers; i.e. crossover fuzzifier & mutation fuzzifier are used as shown in Fig. 4. The inputs to both the fuzzy sets are \u2206Fs1 & \u2206Fs2 respectively i.e. changes in fitness values obtained with GA, as defined by equation 9. Here Fs (g), Fs (g1) & Fs (g-2) denote the fitness function of genetic algorithm during last three generations. Depending upon the changes in the fitness values, FLC decides the changes in the crossover & mutation rate i.e. it makes crossover fraction & mutation rate adaptive [12].\nOutput variable is represented by seven equi-spaced triangular membership functions between -1 & 1. FLC decides the required change in the crossover & mutation rate depending on the change/ rate of change of the fitness function during last three generations. More randomness to Genetic Algorithm is induced by randomly adding a few new members after specific number of generations of GA. The complete steps of working of improved GA are given in Fig. 6.\nIn this work, the parameters of PID controller used for quad-copter control i.e. \ufffd\ufffd\ufffd \ufffd\ufffd \ufffd \ufffd\ufffd have been tuned both using canonical as well as improved Genetic Algorithm. For this purpose, a population size of 30 is taken (= 10 times the value of parameter to be optimized) Other parameters of canonical GA have been specified in the table 1."
        ]
    },
    {
        "image_filename": "designv11_30_0003733_icems50442.2020.9290866-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003733_icems50442.2020.9290866-Figure3-1.png",
        "caption": "Fig. 3. Comparison of magnetic pass between non-field coil current and maximum field coil current conditions.",
        "texts": [
            " In the next section, basic theory of the variable flux in the proposed motor is denoted. Magnetic pass of the proposed motor is shown in Fig. 2. From Fig. 2, stator interlinkage flux st generated by permanent magnet and field coil is obtained by st mag r f c mag f f r fmag H L N I RR (1) From Eqn. (1), variable flux range \u03b2 of the proposed motor can be written by 1 11 c mag mag c mag f f r fmag f f r fmag H L R H L N I RR N I R (2) From Eqn. 2, it is effective that maximum magnetomotive force of the field coil is designed for canceling the leakage flux r of magnet. In Fig. 3, the numerical results of the test motor, in which field coil is magnetized for canceling the leakage flux, are shown. From Fig. 3, it can be seen that leakage flux around the magnet, condition of which is non-field coil current, are canceling by magnetizing the field coil. 1) Motor specification In Table II, motor specification is summarized. Core material of stator and rotor is set as same. The stator coil is designed that current density on maximum phase current is not exceed than 25 Arms/mm2. Thanks to this, - 725 - Authorized licensed use limited to: CALIFORNIA INSTITUTE OF TECHNOLOGY. Downloaded on August 13,2021 at 07:27:59 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003708_5.0015728-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003708_5.0015728-Figure4-1.png",
        "caption": "FIGURE 4. Equivalent Elastic Strain Simulation Results",
        "texts": [],
        "surrounding_texts": [
            "The results of the simulation of equivalent elastic strain at the distance between cracks of 1.5 mm have a minimum value of 2.24E-05 mm at the point of 1.5mm while the maximum value at the point of 3.0 mm is 0.10358 mm. Based on figure 6 which has a minimum equivalent elastic strain at a distance between 2.5 mm crack is 2.32E-05 mm at point 2.5 mm while the maximum value is at point 2.5 mm with a value of 0.0687 mm. Based on the strain stress diagram, when the ductile body has passed the yield strength limit, and the stress continues to increase, then the strain will also rise until it reaches the ultimate strength [8]. Objects will experience perfect plastic due to hardening strains, then experience a fracture [9] Based on the results obtained, it shows that the strain that occurs is pressure due to the vertical pressure load, so that the area that has cracks, will have a negative value. However [10], when the pressure load process returns to the opposite direction, the pressure strain will change to tensile strain which results in a maximum value in the fracture area. 040011-3"
        ]
    },
    {
        "image_filename": "designv11_30_0000946_0020-7403(61)90007-8-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000946_0020-7403(61)90007-8-Figure4-1.png",
        "caption": "FIG. 4. (a) A third mode of deformation for a smooth small-angle wedge for a particular configuration. (b) General type of field for third mode",
        "texts": [
            " the value of h/a for t ransi t ion f rom one mode of deformation, Fig. l(a), to the other, Fig. l(b). In Fig. 3 we show for the different wedge-angles how this critical rat io varies wi th ~ for various values of a. I t is found, however, t h a t for a = 15 \u00b0 and ~ = 0 (i.e. \u00a2 = 45 \u00b0) t ransi t ion does not occur even when h/a decreases to its minimum, i.e. when fi = 0 ; the piling-up mode of deformat ion is re ta ined th roughou t the penetra t ion, despite the availabil i ty of the second mode. We m a y int roduce a th i rd mode of deformat ion as shown in Fig. 4(a) for a smooth tool, and for this par t icular case the result for p/2k versus h/a is entered on Fig. 2(a) and denoted by a circle. [A valid hodograph for this mode is shown do t ted in Fig. 4(a).] For the purposes of our calculations, we have assumed a continuous curve, shown by the chain line in Fig. 2(a), to join-up the results for the two operat ive modes. For configurations in which h/a is smaller t han tha t given by Fig. 4(a), we should have to develop the slip-line field as shown in Fig. 4(b). The cutting of metal strips between partly rough knife-edge tools 229 I n Fig. 5, the crit ical ra t io (d /H) agains t ff is p lo t t ed for var ious values of a; d [ H is the f rac t ion of the original s t r ip th ickness p e n e t r a t e d b y the apex of the wedge O, before the plast ic zone pene t ra tes the remain ing thickness of the str ip. of deformation. The dep th of pene t r a t ion d and the length of contac t be tween the wedge tool and the mate r ia l 1 are re la ted b y 3 H c r = d/1 - cos ~ + (h/a)e r sin ~ (4) For ff = 0, the critical values o f d / H for a = 90 \u00b0, 60 \u00b0 and 30 \u00b0 are in ag reemen t wi th those g iven b y Hill 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000617_9781118717806.ch5-Figure5.27-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000617_9781118717806.ch5-Figure5.27-1.png",
        "caption": "Figure 5.27 Three-dimensional plots of Rwn/(\u03c1w lT /\u03b4wb) as functions of h/\u03b4w and n",
        "texts": [
            "61) and Pw1(HF ) = 1 2 Rw1(HF )I 2 Lm = \u03c1wlwI 2 Lm 2\u03b4wb , (5.62) where lT is the mean length of a single turn (MTL) The winding DC resistance of the single turn is RwDC 1 = \u03c1wlT bh . (5.63) From (5.54), the winding resistance of the nth layer normalized with respect to the high-frequency resistance of the first layer Rw1(HF ) is given by Frn = Rwn Rw1(HF ) = Rwn( \u03c1w lT \u03b4wb ) = FRn( h \u03b4w ) = sinh ( 2h \u03b4w ) + sin ( 2h \u03b4w ) cosh ( 2h \u03b4w ) \u2212 cos ( 2h \u03b4w ) + 2(n2 \u2212 n) sinh ( h \u03b4w ) \u2212 sin ( h \u03b4w ) cosh ( h \u03b4w ) + cos ( h \u03b4w ) . (5.64) Figure 5.27 depicts a 3D plot of Rwn/[\u03c1wlT /(\u03b4wb)] as functions of h/\u03b4w and n . Plots of FRn/(h/\u03b4w) as a function of h/\u03b4w for several individual layers are shown in Fig. 5.28. It can be seen that the resistance of each layer FRn/(h/\u03b4w) decreases with h/\u03b4w , reaches a minimum value, then increases, and finally is independent of h/\u03b4w . Figure 5.29 shows these plots in the vicinity of the minimum Layer number n hoptn/\u03b4w Frn(min) 1 \u03c0/2 0.91715 2 0.823767 1.6286 3 0.634444 2.1062 4 0.535375 2.4932 5 0.471858 2.8276 6 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002643_robio.2016.7866657-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002643_robio.2016.7866657-Figure7-1.png",
        "caption": "Fig. 7. Soft-rigid collision Bullet default algorithm.",
        "texts": [
            " Therefore, for simulation purposes, the motion of the clamp has been limited to a plane and the extra DoF has been used to control the closing/opening on different object shapes, such as a sphere and an elongated object shape, that simulate more realistically vessels and, in general, organic tissues. By applying the default collision detection method proposed by Bullet Physics, it follows that the spatula collides with the soft body in a consistent way only for small displacements and slow motion. Fast displacement of the spatula can cause a complete penetration into the soft body surface without further collisions, see Fig. 7 (left). Moreover, the application of only the ray casting method without node association (second step of the algorithm) causes drawbacks if the dimension of the rigid body along the collision direction is not large enough. The collision appears unstable, as in Fig. 7 (right), when both the nodes belonging to opposite surfaces apply collision forces to the soft body even if they are unfeasible. This drawback is overcome by applying also the second step of the algorithm, that detects the unfeasible collisions and excludes them from the study at the current simulation step. The algorithm has been tested with convex rigid objects and with a clamp grabbing deformable objects, see Figs. 8,9,10. The results are realistic both in terms of visual rendering and force feedback, see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure1-1.png",
        "caption": "Fig. 1. A multiple-mode mechanism composed of four antiparallelogram units and four R joints: (a) a multiple-mode mechanism, (b) Its schematic diagram.",
        "texts": [
            " In Section 3 , the kinematic analysis of the mechanism in its two plane-symmetric modes are discussed, and the DOF and the kinematic paths are proposed. Three plane-motion modes derived from three singular configurations of two plane-symmetric modes and the transition between different modes are identified in Section 4 . The above results are verified by using a physical prototype of the multiple-mode mechanism in Section 5 . Section 6 concludes the paper. By using four R joints to connect four identical antiparallelogram units (A j B j C j D j ), a new multiple-mode mechanism is obtained ( Fig. 1 ). This mechanism consists of 16 links and 20 R joints. The axes of the R joints in the same antiparallelogram unit are parallel to each other, and perpendicular to the R joints lying on the side links. The lengths of A j D j and B j C j are denoted by l , and the lengths of A j B j and C j D j are denoted by l 1 , where l 1 > l + d ( d is the diameter of revolute joint). The angle between A j D j and A j B j of the j th antiparallelogram unit are denoted as \u03b8Aj , where j = 1, 2, 3, 4. Actually, the redundant revolute joints along axes of the four R joints do not affect the movement and freedom of the mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure4-1.png",
        "caption": "Fig. 4: Welding torch attitude definition",
        "texts": [
            " The welding torch coordinate system {FT} is shown in Fig. 3. The origin OT is chosen as the endpoint of the welding torch; axis XT coincides with the axis of welding torch and points to the welding seam; axis ZT lies in the plane determined by axis XT and the centerline of the flange, and perpendiculars to axis XT but deviations to the flange centerline; axis YT is determined by the right-hand rule. In order to describe the pose of the welding torch more clearly, the welding angle, the moving angle and the rotation angle are defined as Fig. 4 shows. The welding angle W is the angle between axis XT of {FT} and axis YP of {FP}, and the direction is from YP to -XT. Its value ranges from -180 to 0 degree. The moving angle M is the angle between of axis XT of {FT} and axis XP of {FP}, and the direction is from XP to -XT. Its value ranges from 0 to 180 degree. The rotation angle R is the angle by rotating {FT} around its own axis XT. In this paper the downward or ship welding attitude, which can inhibit the flowing of iron melt well, are adopted to ensure the welding quality"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002540_ssrr.2016.7784286-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002540_ssrr.2016.7784286-Figure1-1.png",
        "caption": "Fig. 1. Assumed shape of the robot and the stairs",
        "texts": [
            " In addition, there are multiple grousers on the periphery of each track to improve friction between the track and the ground. These should be enough to prevent the robot from slipping down. We assume that the rotation counts of the main tracks and the subtracks are obtained by internal sensors, such as rotary encoders, which enables odometry. The robot\u2019s pose is obtained by an IMU that contains acceleration sensors and gyroscopes. The definition of each specification parameter is described in Fig. 1-(i). 978-1-5090-4349-1/16/$31.00 \u00a92016 IEEE 112 In this paper, we deal with an environment that consists of two different horizontal floors that are connected by stairs, with all steps in parallel. The distance (pitch) p and inclination \u03b8s between two adjacent edges of the stairs are constant. We assume that all steps have a horizontal plane, and the existence of a vertical plane is arbitrary. p, \u03b8s, and the number of steps n are known. The width of the stairs is sufficiently larger than that of the robot",
            " 1) Tracked robot: We used the tracked robot Kenaf [6][7] in our laboratory for verification tests (Fig. 9). The Kenaf is a 6-DOF tracked robot: two main tracks for traversal, and four subtracks, which are located on both sides at the front and rear of the robot, and can be controlled independently. However, left and right subtracks are synchronized in the verification tests. Table I indicates the parameters of the Kenaf. Incidentally, the shape of the Kenaf projected from the side is different from Fig. 1. That is because the diameters of the outer pulleys of the subtracks are larger than that of the semicircular parts of the main tracks. However, according to Fig. 9, each Kenaf\u2019s parameter, such as lf , lm, and lr, is equivalent with our proposed model as shown in Table I. In addition, the robot has a 9-axes IMU sensor module RTUSB-9AXIS-00 made by RT Co., Ltd. on the main tracks to detect the pitch and roll angles of the robot with respect to the gravity direction. 2) Mock-up stairs: The pitch p and inclination \u03b8s between two adjacent edges of the stairs are constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002044_tro.2016.2544338-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002044_tro.2016.2544338-Figure14-1.png",
        "caption": "Fig. 14. Rotation of v0 with s\u0302 or r\u0302, in order to reach q. (a) If s\u0302 is fixed, v0 and r\u0302 rotate in opposite directions from x\u0302 toward s\u0302 and \u2212s\u0302, respectively. (b) If r is fixed, s\u0302 and v0 rotate in the same direction toward \u2212r\u0302.",
        "texts": [
            " (46) Proof: Since \u03b1, \u03b2) \u2208 \u03a9\u2212, q lies inside the cone positively spanned by s\u0302 and r\u0302. Thus, cs , cr > 0 by (35). 1) To determine the range for the direction of the initial velocity v0 , we rewrite (32) as v0 = \u2016vr\u2016 ( 2\u2016s0\u2016 7\u2016vr\u2016 s\u0302 + r\u0302 ) where, by (39) and (40), 2\u2016s0\u2016 7\u2016vr\u2016 = cs cr + \u221a c2 s c2 r + 7\u03bcscs 5\u03bcrcr . (47) Thus, v0 has the same direction as( cs cr + \u221a c2 s c2 r + 7\u03bcscs 5\u03bcrcr ) s\u0302 + r\u0302 = \u2016q\u2016 cr x\u0302 + \u221a c2 s c2 r + 7\u03bcscs 5\u03bcrcr s\u0302 (48) where the last step uses (35). This establishes part 1. 2) The sliding direction s\u0302 is fixed. Fig. 14(a) illustrates that r\u0302 and v0 must simultaneously rotate in opposite directions toward \u2212s\u0302 and s\u0302, respectively, for the ball\u2019s final position to stay at q. To see why, we let p be the intersection between the ray from the origin o in the direction r\u0302 and the line incident on q and parallel to s\u0302. A comparison between \u2192 op + \u2192 pq= \u2192 oq and (35) gives us cr = \u2016 \u2192 op \u2016 and cs = \u2016 \u2192 pq \u2016. As \u03b2 \u2192 0\u2212, cs \u2192 0 and cr \u2192 \u2016q\u2016. Thus, cs/cr \u2192 0+ and, consequently, the direction of v0 approaches that of x\u0302. As \u03b2 decreases from 0\u2212 toward \u03b1 \u2212 \u03c0, p moves away from q toward infinity in the direction of \u2212s\u0302. Meanwhile, both cs and cr increase toward \u221e and the ratio cs/cr increases from 0+ toward 1\u2212. The vector in (48), and therefore v0 , rotates counterclockwise from x\u0302 to s\u0302 (both exclusive). 3) The rolling direction r\u0302 is fixed [see Fig. 14(b)]. As \u03b1 increases from 0+ toward \u03b2 + \u03c0, cs and cr increase toward \u221e, and the ratio cr/cs increases from 0+ to 1\u2212. The direction of the vector in (48); hence, that of v0 , approaches s\u0302. Proof of Theorem 5: 1) For convenience, we write Ci for C(\u03b1, \u03b2i) and Li for L(\u03b1, \u03b2i), i = 1, 2. See Fig. 15. Under part 2 of Lemma 11, \u03b22 > \u03b21 implies that v0(\u03b1, \u03b22) is between v0(\u03b1, \u03b21) and x\u0302. Thus, immediately after leaving o, the parabolic segment C2 is between the parabolic segment C1 and the line segment oq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000719_lra.2016.2530854-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000719_lra.2016.2530854-Figure3-1.png",
        "caption": "Fig. 3. Differential Drive Robot Model",
        "texts": [
            " In this section, we present the concepts and definitions used throughout the paper. The kinematic model for a DDR [12] is given by x\u0307p = v cos \u03b8p, y\u0307p = v sin \u03b8p, \u03b8\u0307p = \u03c9 (1) where v and \u03c9 are the translational and angular velocities of the DDR. In practice, we control the individual velocities of the wheels. Therefore, we have that v = r (\u03c91 + \u03c92) 2 , \u03c9 = r (\u03c92 \u2212 \u03c91) 2R (2) where r is the radius of the wheels, and \u03c91 and \u03c92 are their angular velocities. R is the distance between the center of the robot and the wheel\u2019s location (see Fig. 3(a)). For a DDR, assuming vmax > 0, we have that |\u03b8\u0307| = |\u03c9| \u2264 1 R (vmax \u2212 |v|) (3) The angular velocity is inversely proportional to the translation velocity (see Fig. 3(b)). We assume that the values of the translational velocity v and the angular velocity w can be chosen directly as long as they satisfy Eq. (3). As mentioned earlier, we map the convex environment W to the evader\u2019s configuration space simply by removing all points within unit distance from the boundary. We denote this new environment as Q. From now on, we represent the players as points in Q. Now we identify configurations where collisions between the DDR and the evader are possible in W . In Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003413_s10846-020-01249-2-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003413_s10846-020-01249-2-Figure4-1.png",
        "caption": "Fig. 4 3D model of Prototype",
        "texts": [],
        "surrounding_texts": [
            "and optimization of velocities, an overall control framework that includes the end posture closed loop, joint angle closed joop and rope velocity closed lope is designed to achieve the position servo control for the snake manipulator. Finally, the trajectory tracking simulation and experiment are executed and the results verify the control scheme designed based on the velocity mapping.\nDue to the limitation of mass and size, there are limited DOFs for the the traditional joint-series mechanical manipulator, generally less than 7 DOFs, so it cannot operate in narrow space. Therefore, developing a dexterous and lightweight manipulator has become an urgent problem for mechanical designers. Inspired by biological bodies in nature, especially snakes, the researchers designed rope driven snake manipulators.\nThe snake body is the typical representative of dexterous biological torsos with great high flexibility in nature. The spine of the snake composed of numerous vertebras accounts for more than 90% of the body structure, as shown in Fig. 1, which can adapt to the motion in narrow space and provide an important reference for the design of operating robots in narrow and confined environments. There are many similar vertebral segments for the biological spine, which are composed of vertebras and intervertebral discs, as shown in Fig. 2. In Fig. 2, the vertebrae can be regarded as a rigid structure, which supports the whole body, and the intervertebral disc is a flexible structure, which can be deformed and shock-absorbing. On the outside of the spine, muscle groups are distributed to drive the deformation of the vertebral segments.\nThe vertebrae are pulled by the peripheral muscles, and the intervertebral disc acts to connect the vertebrae.\nThe elastic intervertebral disc can transform the kinetic energy to elastic potential energy during the movement to reduce energy consumption. Although the motion between adjacent vertebrae is small, the multiple series of vertebrae are accumulated together to form a large range of motion, including flexion and extension, lateral flexion, and rotation motion, which can generate high flexibility to spine organisms.\nReferring to the biological snake\u2019s spine, the snake manipulator prototype is designed with five segments and 10 DOFs, as shown in Fig. 3. In the figure, 2 DOF motion of each bionic segment is achieved through a universal joint and 3 ropes, where the ropes are applied to imitate the stretching of the biological tendon. The bionic vertebral segment is mainly composed of an upper plate, a lower plate, 3 support columns, and a universal joint. On the upper and lower plates, there are holes for installing ropes. The ropes for driving the vertebral segment are fixed on the lower joint plate, and the ropes for driving the previous segment are guided to the base pose through the mounting",
            "holes. The driving mechanism of the rope is centrally mounted in the base, which can greatly reduce the weight of the arm, as shown in Figs. 4 and 5.\nIn Figs. 4 and 5, the motion range of each universal joint is \u00b140\u25e6, and the arm is 570 mm in length, 80 mm in diameter. The rope with 2.5 mm in diameter is driven by a ball-screw pair. The outer shell of the arm is attached with a two-dimensional code to measure the posture of the manipulator by visual fusion method. Due to the redundancy of the rope compared to the DOFs, inconsistent rope lengths caused by control errors will cause the mechanism getting stuck. Therefore, the flexible cushioning element is designed and installed between the ropes and ball-screw pairs to tolerate the inconsistent rope lengths, as shown in Fig. 6. To meet both the requirements of tolerance and accuracy, the stiffness of the flexible cushioning element has been carefully designed and verified.\nDue to numerous motion joints, To describe and analyze the posture of the snake manipulator prototype, it is necessary\nto establish an improved kinematic model. The prototype is described using the D-H coordinate system, as shown in Fig. 7. In the figure, Cn is the center of the universal joint of the segment n. Pn(n = 1, 2, \u00b7 \u00b7 \u00b7 5) is the position vector of Cn. According to the mechanical configuration, the D-H parameters can be determined, as shown in Table 1.\nIn Fig. 7, the transformation matrix between adjacent D-H coordinates is:\ni\u22121Ti = \u23a1 \u23a2\u23a2\u23a3 ci \u2212c\u03b1isi s\u03b1isi aci si c\u03b1ici \u2212s\u03b1ici aisi 0 s\u03b1i c\u03b1i di\n0 0 0 1\n\u23a4 \u23a5\u23a5\u23a6\nci = cosqi, si = sinqi, c\u03b1 = cos\u03b1i, s\u03b1 = sin\u03b1i\n(1)\nThe end posture can be obtained from (1) is:\nTI = T0 \u00b7 0T1 \u00b7 1T2 \u00b7 \u00b7 \u00b7 I\u22121TI = [ RI pI\n01\u00d73 1\n] (2)\nWhere, I = 10 is the quantity of the D-H coordinates; pI =[ x y z ]T represents the end position vector. RI represents",
            "Table 1 D-H parameters\nLinki 1 2 3 4 5 6 7 8 9 10\nqi \u03b81 \u03b82 \u03b83 \u03b84 \u03b85 \u03b86 \u03b87 \u03b88 \u03b89 \u03b810 di(mm) 0 0 0 0 0 0 0 0 0 0 a(mm) 0 110 0 110 0 110 0 110 0 50 \u03b1(\u25e6) -90 90 -90 90 -90 90 -90 90 -90 90\nthe attitude rotation matrix of the end coordinate system, and can be expressed by the the Euler angles \u03c6, \u03b8, \u03c8 :\nRI = \u23a1 \u23a3 cos \u03b8 cos\u03c6 sin\u03c8sin\u03b8 cos\u03c6 \u2212 cos\u03c8 sin\u03c6 cos \u03b8 sin\u03c6 sin\u03c8sin\u03b8 sin\u03c6 + cos\u03c8 cos\u03c6\n\u2212 sin \u03b8 sin\u03c8 cos \u03b8\ncos\u03c8 sin \u03b8 cos\u03c6 + sin\u03c8 sin\u03c6 cos\u03c8 sin \u03b8 sin\u03c6 \u2212 sin\u03c8 cos\u03c6\ncos\u03c8 cos \u03b8\n\u23a4 \u23a6\n(3)\nThe number of holes is the same as ropes, which is j = 1, 2, 3 . . . 15. The length between the upper and lower plates of the same segment is constant, so the rope length varies mainly depending on the universal joint between the adjust segments. The rope vector d\nj k between the adjust segments\nk and k \u2212 1 is:\nd j k = r j k \u2212 r j k\u22121 (4)\nWhere, k = 1, 2 . . . 5 represents the identifier number\nof segments. r\nj k is the position vector of the hole j on the\nlower plate of segment k. r\nj\nk\u22121 is the position vector of the hole j on the upper plate of the segment k \u2212 1.\nFor the convenience of calculation, the coordinate of the segment k is set as the D-H coordinates numbered 2k by\ndefault. And then r\nj k , r j k\u22121 can be determined as:\nr\nj k = R2k \u03c1 j k + P2k, r j k\u22121 = R2(k\u22121) \u03c1 j k\u22121 + P2(k\u22121) (5)\nWhere, \u03c1\nj k is the position vector of the hole j on the lower\nplate of segment k in the D-H coordinate C2k . \u03c1\nj\nk\u22121 is the position vector of the hole j on the upper plate of segment k \u2212 1 in the D-H coordinate C2(k\u22121). \u03c1 j k and \u03c1 j\nk\u22121 are constant vectors.\nThe length of rope j is determined by d j k and \u03b7 j k in Fig. 8.\nReferring to Eqs. 2\u20135, the length of rope j can\nlj = \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 rem(j,5)\u2211 k=1 \u2225\u2225\u2225d j k \u2225\u2225\u2225 + l j 0 j = 1, 1 + K, 1 + 2K\nrem(j,5)\u2211 k=1 \u2225\u2225\u2225d j k \u2225\u2225\u2225 + rem(j,5)\u2211 k=1 \u2225\u2225\u2225\u03b7 j k\u22121 \u2225\u2225\u2225 + l j 0 j = 2, 3 \u00b7 \u00b7 \u00b7K; K + 2, 8 \u00b7 \u00b7 \u00b7 2K; 2K + 2, 13 \u00b7 \u00b7 \u00b7 3K (6)\nbe calculated, as shown in Eq. 6. The procedure b is executed according to Eq. 2 and the procedure d is executed according to Eq. 6. Therefore, the procedure a is the inverse calculation of Eq. 6 and the procedure c is the inverse calculation of Eq. 2.\nDue to I = 10 > 6 in Eq. 2, the dimension of q is larger than the end pose. Therefore, there is multiple solutions for the procedure c, while the inverse calculation of Eq. 2 is highly non-linear and q can not be expressed with TN (x, y, z, \u03c6, \u03b8, \u03c8) explicitly. To determine the joint angle, the velocity mapping method is utilized as: \u23a7\u23aa\u23a8 \u23aa\u23a9 v = A [ x\u0307 y\u0307 z\u0307 \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 ]T q\u0307 = J+ q v + ( I \u2212 J+ q Jq ) \u03b7\nq = \u222b q\u0307dt + q0\n(7)\nWhere, A = [\nI 0 0 RI\n] ; \u03b7 is a vector in the joint space; q0 is\nthe initial joint angle; Jq is the end-joint Jacobian matrix. From Eq. 6, the nonlinear mapping between l and q can be shown as:\nl = h (q) (8)\nWhere, h (q) is the nonlinear function of joint angle and determined according to Eqs. 2-6."
        ]
    },
    {
        "image_filename": "designv11_30_0002361_1.4034500-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002361_1.4034500-Figure4-1.png",
        "caption": "Fig. 4 New rotor design: (a) 3D CAD model and (b) section view center body (A-A)",
        "texts": [
            " Compared to the former design, it was necessary to re-design the thrust bearing and the displacement device. The displacement device is needed to adjust the axial rotor position. The new setup includes an extended shaft which leads through the thrust bearing. So, a telemetry system can be installed at the rotor shaft end. Rotor Design and Instrumentation. One main aspect for the development of the new rotor system was the capability of interchangeable running surfaces. This requirement is realized using integrated rotor adapters in the form of annular rings. Figure 4(a) shows a 3D-Model of the new rotor system design. The dimensions of all relevant diameters are identical compared to the former solid rotor body concept, in order to enable a direct comparison of future results with previous measurements. The geometrical dimensions of each adapter are identical except for Journal of Engineering for Gas Turbines and Power MARCH 2016, Vol. 139 / 032502-3 Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jetpez/935787/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use the staggered surface at the drive end (flow guides). The shaft and both adapters are made of a standard CrMoV alloy, typically used in steam turbines. The high-loaded screw connections have been designed and calculated according to German VDI 2230 guideline. The inner geometry of the new rotor is depicted by a sectional view in Fig. 4(b), relating to the cutting planes in Fig. 4(a). The instrumentation of thermocouples in the rotating system requires a hollow shaft design in order to rout the thermocouple wiring. Superficial cable guidance is impossible due to the complex rig design. Starting from the open rotor end at the telemetry system, the rotor has a centered bore hole with nearly half of the total rotor length. Two opposing, inclined bore holes build the way out of the system at the drive end (view A-A). Displaced by 90 deg (view B-B), two additional bore holes are implemented to lead instrumentation wires to the displacement end",
            " The four different radial measurement planes are located in a range of 2\u20139.5 mm underneath the rotor surface. The whole development process was supported by FE analyses in order to guarantee a safe test operation. Despite the high thermal and centrifugal loads, material stresses and stability are uncritical for the described rotor design. The new concept includes an interference section between each adapter and the shaft body in order to avoid lifting the adapters from the shaft surface caused by centrifugal forces (c.f. Fig. 4(b)). Figure 7 verifies the uniform radial growth of the combined shaft\u2013adapters system. The appropriate calculations were conducted for a nominal rotor speed of 10,000 rpm and an ambient temperature of 500 C. Brush seal impacts by means of local heat influxes due to frictional contacts have not been considered within the design calculations. In contrast to previous rotor designs [25], centrifugal lifting effects of the adapters, influencing the clearance between rotor and bristle pack, could be excluded"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003771_ssci47803.2020.9308433-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003771_ssci47803.2020.9308433-Figure1-1.png",
        "caption": "Fig. 1: Robot modules",
        "texts": [
            " Here we also discuss the morphological and behavioral traits we track throughout an evolutionary process. Thereafter we present and discuss the experimental results and conclude the paper. Our robot system is based on RoboGen; the bodies of the robots are modular, composed of three types of modules including a joint module actuated by a servo motor [1]. The robot\u2019s brains (controllers) have a network structure where each joint of the body has a coupled oscillator to drive it and where neighboring joints are connected. The robots are composed of three different types of modules: one Core module (Figure 1a), an arbitrary number of Brick modules (Figure 1b) and Joint modules (Figure 1c). The Core module is the \u201chead\u201d of the robot that contains the main logic board and the battery. Only one Core module is allowed for each robot and it has four connection points where other modules can be attached. Brick modules represent the \u201cbackbone\u201d of the robot, in the sense that only through Brick modules the robot can take up arbitrary shapes. Actuation is achieved through the Joint modules. The joints can be attached to another module in two different ways, which differ by 90\u00b0 concerning the axis perpendicular to the attachment plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000763_icengtechnol.2014.7016810-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000763_icengtechnol.2014.7016810-Figure2-1.png",
        "caption": "Figure 2. Principal of quadrotor controJl21",
        "texts": [
            " 1, VTOL is one type of aerial vehicles which has the ability of vertical and stationary flight. Quadrotor UA V is one of the most famous and widely used VTOL vehicles confIgurations [1]. Quadrotor UA V is the target of numerous investigations within the last decade e.g. [2], [3], [5] and [6]. 978-1-4799-5807-8/14/$31.00 @2014 IEEE Quadrotor is commonly defIned as a mechatronic system which has four-rotor in a cross confIguration. The quadrotor is controlled based on the balance between the resultant four forces, shown in Fig. 2, in which two pairs of rotor are rotating in the opposite direction of the other pairs in order to eliminate the aerodynamic torques. Lift of the quadrotor is achieved by increasing (or decreasing) all rotor speeds by the same amount. Increasing (or decreasing) the relative speed of left and right rotors produces torque which leads to rotation around the X-axis by an angle called Rol1( <p). Increasing (or decreasing) the relative speed of the front and rear rotors produces torque which leads to rotation around the Y-axis axis by an angle called Pitch(9)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003598_j.mechmachtheory.2020.104150-Figure15-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003598_j.mechmachtheory.2020.104150-Figure15-1.png",
        "caption": "Fig. 15. Contact pattern for mesh cycle.",
        "texts": [
            " The line of action is always perpendicular to the involutoid curves and hence the entire tooth flank. The intersection between the involutoid hyperboloid and the gear tooth is the involutoid curve. The involutoid hyperboloid becomes planar as the shaft angle approaches zero. The involutoid reduces to the classical gearing involute and the evolutoid is the circular base circle of Fig. 1 . The involutoid hyperboloid is non-degenerate for the special case of helical cylindrical gears. Illustrated in Fig. 15 are screen shots of the contact pattern for the nominal contact. Two scenarios are presented, unmodified tooth flanks and modified tooth flanks. Although qualitative, the modified contact shows reduced addendum-dedendum contact. Graphs of UTE (Unloaded Transmission Error) are provided in Fig. 16 for both unmodified and modified tooth flanks. This third example considers a special case of non-orthogonal axes gearing with a small shaft offset to further illustrate the use of evolutoids and involutoids to define conjugate gear flanks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003618_ever48776.2020.9243048-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003618_ever48776.2020.9243048-Figure6-1.png",
        "caption": "Fig. 6. Rotor eddy current density distribution with different PWM strategies.",
        "texts": [
            " Because if seen from the rotor, the stator slots rotate in the same direction with the 5th (and 11th) harmonic MMF, while for 7th (and 13th) harmonics, the direction is opposite. In general, harmonics with the same amplitude but higher orders will generate more eddy current loss in the rotor. III. SIMULATION RESULTS AND DISCUSSION The fluctuant magnetic flux density in the air-gap induces considerable eddy current in the rotor, which penetrates from the rotor sleeve to the PM and then to the shaft. The eddy current distribution in the rotor is related to the PWM strategies. By way of example, Fig. 6 shows the eddy current distribution of the rotor under four drive conditions. The small red and blue semicircle patterns on the retaining sleeves are a result of slot effect, which will not change with the control strategy. While the massive eddy currents in the PM region are due to the time harmonics of the armature current. As the switching frequency increases, the phase currents become smoother, so the armaturecurrent-induced loss component can be dramatically restrained, hence the rotor eddy current diminishes accordingly. When the carrier ratio is low, each control strategy generates unique current waveform and armature MMF, so the rotor eddy current loss distribution may show great difference. According to the FEM result, the overall eddy current density of the rotor with the HCC PWM is much lower than those with the other PWM Authorized licensed use limited to: Carleton University. Downloaded on June 02,2021 at 06:00:22 UTC from IEEE Xplore. Restrictions apply. control strategies (such as the 7-SVPWM in Fig. 6), at the same average switching frequency of 10 kHz. As the switching frequency goes up, the armature current waveforms under various PWM strategies get similar, so that the rotor eddy current distributions are similar, too. The rotor eddy current loss with the five PWM strategies are plotted in Fig. 7. A dashed line representing for the ideal sine-wave current drive is introduced into the plot, and it stands for the value of the rotor eddy current loss generated by the PM MMF and the fundamental component of armature MMF only"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure16-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure16-1.png",
        "caption": "Fig. 16. The typical configurations of plane-motion Mode IV.",
        "texts": [
            " Since the axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration II ( Fig. 15 ), when the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R2 and J R6 are in the same direction, the joint axes of R joints J R4 and J R8 are in the same direction, but the two groups directions of joint axes are opposite during the movement. Then the mechanism makes a planar motion, which is called plane-motion Mode IV. Its typical configurations are shown in Fig. 16 . When the axes of R joints J R2 and J R8 coincide, the axes of R joints J R4 and J R6 coincide, and the axes of R joints J R3 and J R7 coincide, the configuration is a special configuration of Mode II. Or when the axes of R joints J R2 and J R4 coincide, the axes of R joints J R6 and J R8 coincide, and the axes of R joints J R1 and J R5 coincide, as shown in Fig. 17 , the configuration is also a special configuration of Mode II. Since the mechanism is an open-chain mechanism with 3 links and 2 R joints, the instantaneous DOF of the mechanism under this configuration is 3",
            " 25 , and the outward configurations in Fig. 10 (c) of plane- symmetric Mode II cannot be achieved because of the offset angle. As shown in Fig. 26 , from the bifurcation configuration in Fig. 26 (a) (which corresponds to configuration (6) in Fig. 22 ), plane-symmetric Mode I ( Fig. 26 (b) and (c)) and plane- symmetric Mode II ( Fig. 26 (d) and (e)) can be obtained. Plane-motion Mode III (corresponding to Fig. 14 ) are shown in Fig. 27 . Figs. 28 and 29 are the snapshots of plane-motion Modes IV (corresponding to Fig. 16 ) and V(corresponding to Fig. 18 ), respectively. Fig. 30 shows the snapshots of the special configurations(corresponding to Fig. 20 ) of this multiple- mode mechanism. Similar to Bricard-like mechanism [61] , the multiple-mode mechanism discussed above can also be used as a construction unit to construct deployable mechanisms. As shown in Fig. 31 , a quadrangular prism deployable mechanism composed of two identical multiple-mode mechanisms connected by eight spherical joints is constructed. During its movement, the two identical multiple-mode mechanisms are mirror-symmetrical"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000712_itsc.2015.311-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000712_itsc.2015.311-Figure11-1.png",
        "caption": "Fig. 11. a) Car-like robot Manta; b) Joints set of car-like robot",
        "texts": [
            " Under most situations, the robot is capable of avoiding the obstacles inside the GPC, but at times the GPC can be totally blocked. Fig. 10(a) shows that there are two obstacles on the way blocking the entire GPC. In such casse, the robot is unable to avoid the obstacles inside the original GPC, and the global path planner needs to be restarted to generate a new global path based on the updated environment. Fig. 10(b) shows that the normative path and the GPC are reproduced based on the updated environment. The GPC-based path planning is tested with the simulated car-like robot, Manta, (Fig. 11) in a small virtual city (Fig. 12(a)) under robotic simulator V-REP. The simulation is programmed with c++ and runs on a computer with an Intel Core Duo(TM)2 CPU P8700 2.53 GHz and 4GB RAM under a 64 bit Windows 7 operating system. Fig. 12(b) shows the 750\u00d7 600 grid map of the virtual city. The GVD of the grid map is extracted with VPT [19], and the computation time is 489.21ms. The simulated human figures are randomly set and moving around dynamically. The car-like robot must follow the global path to reach the goal and dynamically avoid the human figures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure11-1.png",
        "caption": "Fig. 11. The MRR of the math model.",
        "texts": [
            " In the finish process, the highmargin of MRR Please cite this article as: S.-H. Chen, Z.-H. Fong, Study on the cutting time of the hypoid gear tooth flank, Mech. Mach. Theory (2014), http://dx.doi.org/10.1016/j.mechmachtheory.2014.01.012 Fig. 12. Instantaneous TCMD of the cutting-spindle motor. is 7 mm3, and the maximum roll rate is 9\u00b0/s. The roll rates of the CMRR cutting plan are shown in Fig. 9(b), and the corresponding MRR and cycle time are shown in Fig. 10(a) and (b), respectively. The simulated MRR with respect to time is shown in Fig. 11. The high margin of MRR for roughing (h1) is 18 mm3/s, and the duration time (t1) is 4.947 s. The high margin of MRR for finishing (h2) is 7 mm3/s, and the duration time (t2) is 2.015 s. The ratio mh\u2032 of h1 to h2 is 2.57, and the ratio mt\u2032 of t1 to t2 is 2.455. Fig. 12 shows the instantaneous TCMD of the cutting-spindle motor recorded by the FANUC servo-guide. Take the average per 1 s of the TCMD data, and the result is shown in Fig. 13. The maximum TCMD for roughing (h3) is 0.563%, and the duration time (t3) is 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.80-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.80-1.png",
        "caption": "Figure 3.80 Encoder assembly",
        "texts": [
            "\u201d Therefore, basic knowledge of the characteristics of various types of feedback devices is necessary in order to select the type best suited to meet the performance specifications of a closed loop motion control system. The following sections describe the characteristics of various feedback devices, emphasizing optical encoders, resolvers and potentiometers. Optical encoders transform rotary or linear motion into electronic signals for processing by the system controller. In an optical encoder a light source projects through a focusing graticule onto a masked photo-diode/transistor (see Figure 3.80). System Components 147 The light beam is sequentially interrupted by transparent and opaque areas on a rotating disc, resulting in a near sine wave signal. Depending on the design of the disc and associated optics, encoders provide either incremental or absolute shaft angle information. By interpreting the frequency or period and the phasing of these signals, the controller can determine the incremental or absolute position, and the velocity, acceleration and direction of the system motion. In addition, due to the advent of the brushless motor, encoders now also provide three signals, U, V and W to be used by the motor controller to implement the commutation of the motor stator windings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000006_ls.1481-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000006_ls.1481-Figure3-1.png",
        "caption": "FIGURE 3 Schematic diagram of gas face seal textured with generalised ellipse dimples",
        "texts": [
            " Generalised ellipse is presented as a conventional ellipse (n = 2, \u03b1 = 180\u00b0), a triangle (n = 3, \u03b1 = 120\u00b0), and a pentagon (n = 5, \u03b1 = 72\u00b0) at different values of unit number, respectively, which can be called as polygonal generalised ellipse. However, generalised ellipse is presented as heart\u2010shape (\u03b1 = 0\u201090\u00b0), boomerang\u2010shape (\u03b1 = 90\u2010180\u00b0), and conventional ellipse (\u03b1 = 180\u00b0) with the increase of included angle between two semi\u2010major axis. A new type of generalised ellipse dimple gas face seal is proposed by introducing generalised ellipse to characterise the boundary of surface dimple. A schematic drawing of a gas face seal with generalised ellipse dimples is presented in Figure 3. The textured region adjacent to the high\u2010pressure boundary at outer radius, do, extends from dp to do. The untextured dam region locates at the seal inner radius from di to dp, the film thickness where is h0, as shown in Figure 3A,B. To reduce computational time, a dimple column with circumferential angle 2\u03c0/n\u03b8 is selected as a computing period, as shown in Figure 3 C. Two typical generalised ellipse dimple elements with a semi\u2010major axis a, semi\u2010minor axis b, dimple depth hp, included angle of semi\u2010major axis \u03b1, and inclined angle \u03b8 are presented in Figure 3D,E. For the purpose of considering the effect of dimple profile parameter, short axis ratio B and long axis ratio A are defined as follows: B \u00bc b a ; (1) A \u00bc nra ro \u2212 rp ; (2) where ro and rp are the outer radius of seal ring and boundary radius between textured portion and sealing dam, respectively, and nr is the dimple column number in the radial direction. The sealed gas is assumed to be compressible and with a constant viscosity \u03bc. The generalised Reynolds equation for compressible fluid can be expressed by \u2202 r\u2202\u03b8 h3 \u03bc \u2202p2 r\u2202\u03b8 \u00fe \u2202 \u2202r h3 \u03bc \u2202p2 \u2202r \u00bc 12\u03c9 \u2202 ph\u00f0 \u00de \u2202\u03b8 : (3) For the generalised ellipse dimple surface, the local film thickness, h(r,\u03b8), is just the film thickness h0 in untextured region, while it is the sum of film thickness h0 and dimple depth hp in textured region"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003333_rpj-01-2020-0016-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003333_rpj-01-2020-0016-Figure2-1.png",
        "caption": "Figure 2 Schematic diagram of GMA-based AM process for the inclined component",
        "texts": [
            " The chemical compositions of the substrate and wire are listed in Table 1. Figure 1(b) displays the schematic and dimensions of the component. The substrate is 200mm in Positional wire and arc additive manufacturing Dongqing Yang, Jun Xiong and Rong Li Rapid Prototyping Journal length, 110mm in width and 12mm in height. The inclined wall has a wall length 150mm and a wall thickness varying between 7 and 7.8mm. Except the bottom part, the component is tilted at 45\u00b0 with the horizontal substrate. During the fabrication process, a positional deposition method shown in Figure 2 is adopted, in which the torch axis is set to be parallel to theXY plane with an angle to the y-axis of 45\u00b0. This positional fabrication technique can greatly reduce the movement of components and get rid of the dependence on the positioner. The main process parameters used in the experiment were as follows. The arc voltage was 19V and the deposition current was 150A, and the torch travel speed was 5mm/s. 95% Ar1 5% CO2 at a flow rate of 18L/min was used as the shielding gas. After each layer was deposited, an interval time of 42 s was applied to obtain a proper inter-layer temperature",
            "What is more, the distribution parameters of the heat source are unchangeable during the whole modeling process. Nevertheless, these parameters are variable in the actual experiment. It can be concluded that the calculated thermal cycle curves coincide with the measured data, validating the accuracy of the proposed simulationmodel. Figure 7 shows the molten pool size as the heat source is located at the midpoint of the third layer. As represented in Figure 7(a), the molten pool captured by the digital camera shown in Figure 2 has a length of 14.8mm. Here, the molten pool length in front of the electrode is ignored because it is difficult to measure and only accounts for a very small proportion of the whole length. The actual dimensions of the molten pool are used to adjust the distribution parameters of the heat source used in the simulation process. Figure 7(b) gives the simulated results, in which the molten pool length is Positional wire and arc additive manufacturing Dongqing Yang, Jun Xiong and Rong Li Rapid Prototyping Journal 14"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000001_demped.2019.8864908-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000001_demped.2019.8864908-Figure6-1.png",
        "caption": "Fig. 6: 2-D Finite Element Model",
        "texts": [
            " The machine is operated at the same operating conditions as it is operated for commanded voltages method. FEA simulations were carried out on a three-phase 8-pole distributed winding interior permanent magnet synchronous machine. The machine was simulated under three conditions: healthy, 25% eccentricity and 50% eccentricity in order to predict experimental results. The machines specifications are given in table I. ANSYS Maxwell was used to simulate the machine under 2-D FEA (Finite element analysis). The FEA model is shown in Fig.6. Static eccentricity fault was implemented by shifting the stator in downward direction (negative y direction). The amount of shift is proportional to the severity of the fault to be emulated. The machine is controlled using Field Oriented Control (FOC) scheme using LABVIEW Real Time. The scheme is shown in Fig.7. The tested machine was operated at constant torque, whereas the dynamometer was operated at constant speed. Fig.8 shows the experimental set up that indicates tested machine and dynamometer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure7-1.png",
        "caption": "Fig. 7. Diagram of the IEHSD calculation model.",
        "texts": [
            " To solve the inverse kinematics problem is determined: - The elongation of hydraulic axis: ,i i i= + -L p r R (3) where ri \u2013 Vector coordinates of B iB in the moving frame (B, X, Y, Z), Ri \u2013 Vector coordinates of A iA in the reference frame (A, X0, Y0, Z0), p \u2013 Vector coordinates from point B in the moving frame to point A in the reference frame, [ ] ,T p p px y z=p (4) - The velocity of hydraulic axis: ,T T i i i iL = \u00d7 =u L u v& & (5) where ui is unit vector, =v p& is velocity vector: [ ] ,T x y zv v v=v (6) - The acceleration of hydraulic axis: ,T T T i i i i iL = \u00d7 = =u L u v u a&& && & (7) where a is acceleration vector: [ ] .T x y za a a=a (8) The dynamics model of TPM with IEHSDs can be de- scribed as follows: ( ) ( , ) ( ) ( ) ,T r C G i L+ + + =Mq F q F q q F q J F L&& & & (9) where M \u2013 Load mass matrix, ( )rF q& \u2013 Forces due to friction, ( , )CF q q& \u2013 Forces arising from centrifugal and Coriolis forces, ( )GF q \u2013 Force of gravitation effects, T iJ \u2013 The force Jacobian matrix, ( )LF L \u2013 Load force, 1 2 3( ) [ ] .T L F F F=F L (10) A diagram of the IEHSD calculation model is shown in Fig. 7. The relationship between the elongation Li and the coordi- nate yi of the hydraulic axis, therefore, is: 0 ,i i iy y L= + (11) where y0i \u2013 Initial position of the actuator piston. The general dynamic model of IEHSD was written as follows: ( ) ( ) 1 1 2 2 1 1 1 1 2 10 1 2 2 2 1 2 20 2 2 1 ( , ) ( ) 1 ( ) ( ) i y i fi i i i y L i Gi i i i Z i Z i e Z i v i l i i i i v i l i i i le i y b y F y p A p m m A p F sign y F z b z c z F m Kp q A y k p p V A y t Kp q A y k p p V A y t k p \u00ec \u00e9= - - + +\u00ef \u00eb+\u00ef \u00ef - - - \u00f9\u00fb\u00ef \u00ef = - - +\u00e9 \u00f9\u00ef \u00eb \u00fb \u00ef \u00ed \u00ef \u00e9 \u00f9= - - -\u00eb \u00fb\u00ef + \u00ef \u00ef \u00e9= - + + - +\u00eb\u00ef -\u00ef \u00ef - \u00f9\u00fb\u00ee && & & && & & & & & (12) where yi \u2013 Displacement of actuator piston, zi, \u2013 Displacement of control valve spool, pi \u2013 Pressure, my \u2013 Mass of the piston and piston rod, mL \u2013 Load mass, mz \u2013 Mass of valve spool, by, bz \u2013 Coefficients of viscous friction, cz \u2013 Coefficient of spring rate, A1, A2 \u2013 Piston surface, V10, V20 \u2013 Initial stroke volume, FGi \u2013 Gravitational force, Fi \u2013 External forces, K \u2013 Bulk modulus, kl \u2013 Coefficient of internal leakage, kle \u2013 Coefficient of external leakage, Fe \u2013 Electromagnetic force in the solenoid valve, \u00b7Ff \u2013 Friction force in the hydraulic cylinder seals, ( ) ( ) ( ) ( ) 1 2 exp ( )iy t fi Ci si i s i i F F F sign y t k p p s-= + + + + & & (13) where FCi \u2013 Coulomb friction force, Fsi \u2013 Static friction force, s \u2013 Variable parameter, ks \u2013 Constant coefficient, ks = (ms p ds ls)/2, ms \u2013 Coefficient of friction in seals, ds \u2013 Seals diameter, ls \u2013 Seal length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001377_2013-01-1898-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001377_2013-01-1898-Figure3-1.png",
        "caption": "Figure 3. Finite element model of the gearbox with the full FE mesh for the housing (Cut view).",
        "texts": [],
        "surrounding_texts": [
            "The time-domain dynamic analysis on the gearbox is performed using the finite element-contact solver especially suited for gear dynamics [3,4,5]. Finite Element Contact Mechanics Model for Gear Dynamics The only inputs for the finite element contact mechanics model for dynamics are the input torque and speed along with the geometry and material properties of various components of the gear box. There is no need to externally specify the time-varying stiffness or the static transmission error as the excitation to the dynamics model as required for lumped parameter models. These are all intrinsically modeled in the finite element model. The solver uses a combined surface integral and finite element solution approach to model the contact between gears efficiently [2]. Detailed description on the finite element-contact mechanics formulation for 2D gear dynamics is provided in [3]. An extension to the above formulation for 3D gear dynamics is presented in [7]. The equations of motion for gear box assembly are given by (1) where xf represents the finite element deflections with respect to the reference frame attached to rotating gear bodies, xr is the small rigid body deflections of the reference frame superposed over the large kinematic trajectories of the gears. Rayleigh damping model is used to add damping to gears, shafts and housing, Cff = \u03b1Mff + \u03b2Kff. Viscous damping is added through the bearing damping matrices. The helical gear pair geometry used in this study is adapted from the spur gear pair of involute contact ratio 1.37 from [6], with the addition of the 30 deg. helix angle. Table (1) shows various parameters of the helical gear pair. The shafts, bearings and the housing are arbitrarily picked to build the entire gear box. Figure (3) displays the finite element model with full FE mesh for the housing. Housing Models Two different housing models are considered for this study. In the first case a full FE mesh for the housing is used, whereas in the second case the full FE mesh is replaced by a reduced model of housing with condensed matrices. The housing FE mesh chiefly consists of shell elements, except the bearing sleeves with solid elements. The reduced housing model is obtained by employing the \u201cCraig-Bampton\u201d or \u201cFixed-Interface\u201d CMS method using a commercial FE software package. CMS methods are used to reduce the size of FE matrices to a smaller size of boundary degrees of freedom and truncated set of normalized mode generalized coordinates [10]. To use Craig-Bampton method we first identify the boundary nodes on the housing FE mesh that are going to interface with the rest of the gear box model. In our case, the four constraint nodes and the nodes that are going to connect to bearings are the boundary nodes. To reduce the number of boundary nodes attached to bearings we created a web of rigid elements connecting boundary nodes to a single node at the center of the bearing sleeve. Therefore we will eventually have four boundary nodes for the four bearing interfaces and four more boundary nodes for the constraints. Figure (4) shows the housing finite element mesh used in the CMS method. We also choose the number of normal modes to truncate the modal basis. Based on these two criterions the finite element software package computes the condensed mass and stiffness matrices using Craig-Bampton method. The equations of motion for a FE model may be written as, or, (2) where the degrees of freedom associated with the boundary nodes are represented by B and the interior nodes by I. The physical coordinates of the housing xH may be represented by the component generalized coordinates xh which is a vector of boundary degrees of freedom xB and a set of truncated generalized modal coordinates \u03b7 as follows, (3) where T is the transformation matrix. Equation (3) is important for back-computing the physical coordinates once the component generalized coordinates are known. The condensed stiffness and mass matrices are given by (4) In Craig-Bampton method, the transformation matrix T is computed from the redundant static constraint modes, SIB, and the fixed-interface normal modes, \u03a6I (these are the eigen vectors computed with interface nodes fixed) [10]. (5) where, and I is an identity matrix. Dynamic Analysis Results & Discussion Time-domain dynamic analysis is performed at gear speed 2100 RPM or 35 Hz. The gear mesh frequency at this speed is 1750 Hz. The bearing reactions for the two models with different housing models are compared in Figures (5) and (6). The transients are excited due to the step input of torque. At time zero the torque applied on the system is zero and at the next time step operating torque is applied. Sufficient number of time steps is chosen for the system to reach steady state. The time step is chosen small enough to capture frequencies up to the 5th mesh frequency harmonic. However we deal with only the frequencies within the first two mesh frequency harmonics in this study. Figure (5) shows the transient bearing reactions from the two cases with different housing models. Both the time and frequency domain comparisons match well. The discrepancy in the frequency spectrum for frequencies above 4000 Hz could be attributed to the truncation of normal modes while building the reduced housing models. The steady state bearing reactions from both the housing models are shown in Figure (6). The time domain and frequency domain comparisons match very well. As expected we find gear mesh frequency and its higher harmonics in the frequency spectrum of the steady state response. The first two mesh frequency harmonics at f1=1750 Hz, and f2=3500 Hz are displayed. Thus we establish that the reduced housing model may be effectively used in place of the full FE mesh of the housing. From this point forward we will use the dynamics results from the analysis with the reduced housing model. STEADY STATE ACOUSTIC ANALYSIS A frequency domain boundary element solver [8, 9] is used to estimate the air-borne noise radiated by the housing under steady state conditions. The steady state results from the above time-domain dynamic analysis are transformed to the frequency domain to be used for the acoustic analysis. Housing Frequency Response Function (FRF) We require the housing frequency response function (FRF) to solve the acoustic radiation problem. Below two different methods are detailed for computing the housing FRF. Method A: Housing FRF from component modes superposition In this method we utilize the generalized coordinates of the reduced housing model and the component modes to compute housing FRF. First, the steady state generalized coordinates of the reduced housing model are transformed to the frequency domain. Then, we perform modal superposition using the component modes and generalized coordinates. From equation (3) we obtain the response on the entire housing FE mesh. Please note that we need to have the component mode matrix to use this method. Method B: Housing FRF from forced response analysis using bearing loads In this method we use bearing loads to perform forced response analysis on the full FE mesh of the housing. The steady state bearing loads from the time-domain analysis above are transformed to the frequency domain and are applied as forced in housing forced response analysis. Boundary Element Acoustic Model Boundary element methods are well suited to solve exterior acoustic radiation problems. The main advantage comes from the requirement of discretizing only the surface of the vibrating structure and not the acoustic domain itself. We employ a powerful BEM solver accelerated by fast multipole method (FMM) [8]. Acoustic Analysis Results & Discussion The acoustic analysis is performed only at the frequencies related to the first (f1=1750 Hz) and second (f2=3500Hz) mesh frequency harmonics as the steady state response is dominated by these frequency components. Figure (7) shows the sound pressure level contours over the surface of the housing and over a spherical field mesh at the frequency 1750 Hz. Figure (8) compares the radiated sound power computed from methods A and B. The differences between the two methods may be attributed to the truncation of modes in method A as compared to method B."
        ]
    },
    {
        "image_filename": "designv11_30_0002846_1077546320909968-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002846_1077546320909968-Figure1-1.png",
        "caption": "Figure 1. Multiple autonomous substructure models of the flexible beam.",
        "texts": [
            " According to the above results, we will build the equivalent model of the large flexible structure, which can reflect the flexible effect and the nonlinear characteristics of the structure due to interior coupling. In this section, we will propose the multiple autonomous substructure modeling method of the flexible structure, and each autonomous substructure equips its own sensor, actuator, localized controller, and a means of communicating with their neighbor substructures. Consider an undamped flexible cantilever beam shown in Figure 1 with N actuators and sensors.Ox,y,z is the inertial frame of reference with the origin O, which is at the crossover point of the central axis and root of the flexible beam. The x axis is along the extension direction of the flexible cantilever beam, y axis is perpendicular to the beam\u2019s upper surface, and z = x \u00d7 y. The flexible beam just has bending vibration in plane (x, y), and the elastic deformation is supposed to be small. It is supposed that the beam is discretized as N autonomous substructures as shown in Figure 1, and each autonomous substructure can perform the control tasks without external interference with own sensors, actuators, and controller. The undamped flexible cantilever beam is assumed to be divided into N autonomous substructures. Each substructure is regarded as a rigid body segment, and the rigid segments are connected by smooth hinge and bending spring. Therefore, we can obtain the multisubstructure discretization model of the undamped flexible cantilever beam as shown in Figure 2. As shown in Figure 2, i is the number of the ith substructure, (i) is the number of the equivalent spring, and ki is the bending coefficient of the equivalent spring (i) according to Wang and Huston (1994), and ki \u00bc 2EiEi\u00fe1IiIi\u00fe1 EiIili\u00fe1 \u00fe Ei\u00fe1Ii\u00fe1li where li and Ii are the length and moment of inertia of the ith autonomous substructure, respectively, and Ei is the elastic modulus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000408_1.4040999-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000408_1.4040999-Figure2-1.png",
        "caption": "Fig. 2. Blade Mode Families 1F, 2T and Mixed of Simplified Compressor Blisk",
        "texts": [
            " The first blade-dominated mode families are highlighted with a solid line and black markers. In addition, grey markers identify modes with relevant elastic strain energy ratios in blade and disk. In this paper, the identification of mistuning pattern is focused on the first blade-dominated modes outlined in Fig. 3 as first flap (1F), two basic torsion modes (1T/ and 1T\\), second torsion (2T) and a mixed blade dominated mode. The modeshapes of first flap (1F), second torsion (2T) and a mixed mode are shown in Fig. 2. (a) (b) Fig. 4. Industrial High-SPeed Centrifugal Compressor (a) and Sector Model Equivalent analyses are performed with the second object of investigation, shown in Fig. 4. In contrast to the nodal diameter plot of the axial blisk, there is a splitter blade dominated mode family in the nodal diameter map highlighted Fig. 5. Blade Mode Families BF1, BF2, BF3 and BF4 nodal diameter (ND) 0 1 2 3 n o rm a liz e d f re q u e n c y [ -] 0.5 1 1.5 2 2.5 3 BF1 BF2 BF3 BF4 Fig. 6. Nodal Diameter Map of Industrial Centrifugal Compressor by a dashed line and blue markers in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000613_icecct.2015.7226205-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000613_icecct.2015.7226205-Figure2-1.png",
        "caption": "Fig. 2. Diffrential drive kinematic model of autonomous mobile robot. a) Autonomous mobile robot\u2019s kinematics model based upon unicycle robot. b) Autonomous mobile robot\u2019s motion model.",
        "texts": [
            " The velocity of the robot can be represented as a linear velocity (forward and backward), and the angular velocity (rotation) of the robot. Given the wheel rotations i.e. right and left, the linear and angular velocity of the deferential drive robot can be computed according to the following equations Eq. 1 and Eq. 2: Hence we can compute the robots pose (x, y, \u03b8) according to the following state space model given in Eq. (4) The autonomous mobile robot is modeled as a simple two-wheel vehicle, as shown in Fig. 2. The non-holonomic kinematic equations of the wheelchair are: x\u0307(t) = v(t)cos(\u03b8(t)) y\u0307(t) = v(t)sin(\u03b8(t)) (1) \u03b8\u0307(t) = \u03c9(t) The kinematic model of the mobile robot has two inputs, the translational velocity, v and the rotational velocity (\u03c9). In discrete time, the mobile robot\u2019s laws of motion are: xk+1 = xk + vkcos(\u03b8k)\u0394t yk+1 = yk + vksin(\u03b8k)\u0394t (2) \u03b8k+1 = \u03b8k + \u03c9k\u0394t The two control inputs, u1 and u2, are generated by processing algorithms applied to the laser signals. The mobile robot\u2019s position update from point (xk, yk) to point (xk+1, yk+1) (Fig. 2) is given by: \u0394S = vk\u0394t = u1Vf\u0394t (3) \u0394\u03b8 = \u03c9k\u0394t = u2Vr\u0394t (4) where Vr and Vf are the maximal rotational and forward velocities, respectively, and \u0394t is the time interval between the two consecutive frames of time. The proposed technique was applied to the robotics simulator running under real-time operating System (RTOS) Ubuntu 8.04 Hardy heron on 2.0 GHz Intel dual core processor having 3 GB of RAM. The environment was created in realtime (distributed under GNU) 2D autonomous mobile robotic software, Player/ Stage as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003852_s00773-015-0322-5-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003852_s00773-015-0322-5-Figure1-1.png",
        "caption": "Fig. 1 The hydrodynamic panel model of the semisubmersible platform",
        "texts": [
            " To solve this problem, the balanced truncation algorithm is used to reduce the order of the state-space model. The HYSY-981 semi-submersible platform is adopted in this study. The main structure of this platform consists of two pontoons, four columns, deck, and derrick. The specifications of the HYSY-981 are presented in Table 1. The frequency dependent hydrodynamic parameters are calculated using the ANSYS AQWA software. The hydrodynamic panel model of the semi-submersible platform underwater geometry is given in Fig. 1. The frequencies are ranged from 0.1 to 3.5 rad/s with 0.025 rad/s steps. To obtain very high frequency radiation damping, 7960 panel elements are chosen. From the radiation damping matrix B0 shown below, it can be seen that there are 10 non-zero parameters because of the symmetrical underwater geometry about the x\u2013z plane. Several typical results of hydrodynamic coefficients are given to describe the state-space model of the convolution integral terms. The radiation damping results of B0 22 and B0 24 in the frequency domain, which are the sway and sway-roll coupled terms, are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.34-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.34-1.png",
        "caption": "Figure 3.34 Stepper motor coil configurations",
        "texts": [],
        "surrounding_texts": [
            "The stator coils in stepper motors exist in a number of different configurations, depending on the type of motor and the use of either a bipolar or unipolar drive circuit. Figures 3.34 and 3.35 show the main type of windings and circuits as follows: \u2022 4 wire: Two phases, two coils; two leads per coil bipolar drive Figure 3.35a Used with permanent magnet and hybrid motors. Requires \u201cH\u201d drive amplifier capable of sequentially reversing coil current to reverse stator field polarity. \u2022 6 wire: Two phases; two coils with each coil center tapped; six leads unipolar drive Figure 3.35b. 84 Electromechanical Motion Systems: Design and Simulation Used with permanent magnet and hybrid motors. The center taps are connected to the supply voltage and the coil end wires are sequentially connected to ground by the amplifier to reverse stator field polarity (In a variation, the center taps are connected internally and brought out as a single lead \u2013 5 wire) \u2022 8 wire: Two phases; two coils per phase; eight leads. Unipolar drive/bipolar drive (Figures 3.35(a) (b)) Used with permanent magnet and hybrid motors. With the coils in each phase in series and the common connected to the supply voltage, a 6 wire configuration is created. With the coils in each phase in series and the common not terminated, a 4 wire high voltage/low current configuration is created. With the coils in each phase connected in parallel, a 4 wire low voltage/high current configuration is created. \u2022 8 wire: Four phases; one coil per phase; eight leads. Unipolar drive (Figure 3.35c) Typical four phase variable reluctance configuration One end of each coil is connected to the supply voltage The opposite end of each coil is sequentially connected to ground by the amplifier to provide field excitation. Note: There is no universal color coding/polarity identification among motor suppliers for the lead wires. Consult each supplier\u2019s data sheet for the connections required for proper operation."
        ]
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure5-1.png",
        "caption": "Fig. 5: Ideal welding attitudes",
        "texts": [
            " The welding angle W is the angle between axis XT of {FT} and axis YP of {FP}, and the direction is from YP to -XT. Its value ranges from -180 to 0 degree. The moving angle M is the angle between of axis XT of {FT} and axis XP of {FP}, and the direction is from XP to -XT. Its value ranges from 0 to 180 degree. The rotation angle R is the angle by rotating {FT} around its own axis XT. In this paper the downward or ship welding attitude, which can inhibit the flowing of iron melt well, are adopted to ensure the welding quality. These two ideal welding attitudes are drawn in Fig. 5. Two basic assumptions are made here. The relative pose transformation between the tool and the workpiece is not considered here. Instead the transformations between tools of multiple robots are taken into account via cooperation constraints. Moreover, considering tools are always fixed to the end of flange, the kinematic chain of multi-robot system composed by a master robot and several slave robots can be simplified as Fig. 6 shows. For the master robot, the kinematic chain from the master robot base to the master robot tool is considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001238_isco.2013.6481129-FigureI-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001238_isco.2013.6481129-FigureI-1.png",
        "caption": "FIG. I: TORA SYSTEM",
        "texts": [],
        "surrounding_texts": [
            "In this section, a generalized control law based on block backstepping has been proposed to address the stabilization problem of underactuated mechanical system. The steps of the control algorithm design are as follows: Step 1: Define the new control variable z\\ as follows z,=q,-k(qz+g, Pz-gzp,) . . . (4) where, k is a design constant. The dynamics of z\\ is expressed in equation (3) z, = p,-k(pz +g,P2 +g,p, -g2P,-g2P,) = p, -k(p, + g, (I, + g2U)+ P2dg,.Q-g, (J; + g,u)-p,dg,.Q) = p, -k (p, + gJ2 + p,dg,.P -g2J; -p,dg2 .p) . . . (5) In above representation Q = [q, q, r ,P = [PI P2 r dg = [\ufffd \ufffd] and for ease of representation J.. ( q, p), 8q, 8q, J;(q,p) , g, (q), g2(q)are represented by J..,J;,g\"g2 respectively. Step 2: Choice of stabilizing function for z\\ subsystem. A conventional choice for stabilizing function is given below, a, = -c,z, +k(P2 + g,J; + P2dg,.P-g2J.. -p,dg2\u00b7P ) Step 3: the second error variable is defined as follows z2=p,-a, ... (6) with the above definition of second error variable, the dynamics of first error variable becomes: . . . (7) Dynamics of second error variable Z2 can be expressed as i2 = p, -eX, =J.. + g,u + c,i, -k[{P2 + fsJ2 + gJ; + P2dg,.p = h +g\\u+q (Z2 -qz\\) -k[{.I; + g2u+dg\\ .p.l; + g\\%.X +(.1; + g2u)dg\\.p + P2 (ag\\ (h + g\\u) + 8g\\ (.I; + g2U)J 8q1 8q2"
        ]
    },
    {
        "image_filename": "designv11_30_0000844_1.4027398-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000844_1.4027398-Figure6-1.png",
        "caption": "Fig. 6 Schematic diagram of the longitudinal section of the bearing",
        "texts": [
            " (1): \u00bd\u00f0Rb h1\u00de sin a1 \u00fe x 2 \u00fe \u00bd\u00f0Rb h1\u00de cos a1 \u00fe z 2 \u00bc R2 j \u00bd\u00f0Rb h2\u00de sin a2 x 2 \u00fe \u00bd\u00f0Rb h2\u00de cos a2 \u00fe z 2 \u00bc R2 j (1) where Rb is the bearing radius, Rj is the journal radius, and a1 and a2 are the angles between the axis of two eddy current sensors and the negative direction of the z coordinate axis, respectively. Journal of Tribology JULY 2014, Vol. 136 / 031708-3 Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/930316/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The longitudinal section of the bearing is shown in Fig. 6. The location of the journal center at the axial direction of the bearing is obtained by the measured distance s between the front end plane of the crankshaft and the head end plane of the eddy current gap sensor. 4.1. 3D Orbit of the Journal Center of the Crankshaft Bearing. The 3D orbit of the journal center of the No. 1 main bearing and its projection in three coordinate planes in one operating cycle of the engine at 3200 rpm and full load are shown in Figs. 7 and 8. It is shown that the orbit of the journal center of the bearing is a 3D spatial curve because there is movement of the crankshaft journal along the axial direction of the bearing in engine operation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000774_cjme.2013.03.573-Figure6-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000774_cjme.2013.03.573-Figure6-1.png",
        "caption": "Fig. 6. Inner and outer cone surfaces",
        "texts": [
            " Auxiliary coordinate systems Sa1 and Sb1 are used to help the installation of the pinion with respect to the cradle and the head cutter. As can be seen from Fig. 5, there are five parameters in the machine-tool settings for the generation of the pinion tooth surfaces. They are the installment angle q1, the radial setting Sr1, the vertical offset Em1, the sliding base XB1, and the machine center to back XD1. They are determined by the local synthesis according to the meshing performances at the mean contact point. When the head cutter revolves about its own axis, the straight blade formed a cone as in Fig. 6, which acts as the tooth surface of the cutter. The equations for the cone surface and its unit normal in the coordinate system Sp are represented as follows: p p p p p p p p p p p p p ( sin )cos ( , ) ( sin )sin , cos R u u R u u \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1             r (5) p p p p p p p cos cos ( ) cos sin , sin \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1             n (6) where up and \u03b8p are the tooth surface parameters; \u03b1p is the pressure angle; Rp is the point radius of the head cutter. The upper sign in notations \u201c \u201d and \u201c  \u201d is used to the outer cone or the concave side of the pinion, while the lower sign is used to the inner cone or the convex side of the pinion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003141_j.ijsolstr.2020.06.016-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003141_j.ijsolstr.2020.06.016-Figure2-1.png",
        "caption": "Fig. 2. (a) Parameters for the Miura-ori unit cell; mountain and valley creases are indicated by the black solid and black dashed lines, respectively; (b) partially folded configuration of the unit cell, the fold angle of valley crease is positive, i.e, q4 > 0, whilst those of mountain creases are negative; (c) crease pattern for the Miura-ori fold consisting of 3 3 unit cells with the parameters a \u00bc b \u00bc 1 and a \u00bc p=3; (d) dimensions of the folded form. In (c) and (d), the fold angle of the red crease is controlled with the Lagrange multiplier. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            "2 illustrates the sequential folding of the origami crane, a typical origami example of multiple DOFs, using three folding substeps, where different sets of creases are successively controlled with the Lagrange multiplier. The last two examples in subSections 5.3 and 5.4, folding simulations of Waterbomb and Waterbomb tessellation, respectively, show that the equilibrium configuration can be reached iteratively using Algorithm 2, when the rigid origami is equipped with rotational springs at its creases. The Miura-ori unit is comprised of four identical parallelograms characterized by the parameters a; b, and a (Fig. 2(a) and (b)). The folding motion of a Miura-ori fold with 3 3 unit cells is simulated by controlling the angle q1 as shown in Fig. 2(c). For the planar Fig. 3. (a) Fold angle q2 versus q1; the insets are frames forq1 \u00bc 30 ; 90 and 175 ; ( the results obtained from the simulation and the dashed lines are the analytical solutio state with all fold angles equal to 0, the third equation in Eq. (7) will degenerate as ci, the (2,1)-entry of the derivative matrix in Eq. (3), equals 0 (Tachi, 2012). This reflects the fact that several folded forms are permissible as the mountain and valley crease assignment is missing. In the simulation, the initial fold angles are prescribed with small values, for instance 1 , whilst the positive and negative signs are assigned for the valley and mountain fold angles, respectively. The numerical error can be eliminated by the iterations in Algorithm 1. During the simulation, the controlled angle q1 (see the red crease in Fig. 2(c) and (d)) decreased from 0 to 180 by 5 in every folding step. Fig. 3(a) shows q2 versus q1, which is in agreement with the analytical solution (the dashed line): tan q2=2\u00f0 \u00de \u00bc cos a\u00f0 \u00de tan q1=2\u00f0 \u00de (Evans et al., 2015). The insets in Fig. 3(a) are the folded form, with q1 \u00bc 30 ; 90 , and 175 . The dimensions of the folded form (Fig. 2(d)) are measured at the end of each folding step. Fig. 3(b) shows the change in width and length of the 3 3 Miura-ori fold during the folding process. In fact, the length and width agree exactly with the pertinent analytical predictions, as the numerical error is negligible at the end of each folding step. The in-plane Poisson\u2019s ratio, defined as mLW \u00bc dL L = dW W (Wei et al., 2013), is calculated numerically as mhLW \u00bc Li\u00fe1 Li Li = Wi\u00fe1 Wi Wi , where the subscript i and i\u00fe 1 correspond to the number of folding steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001978_tmag.2014.2334312-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001978_tmag.2014.2334312-Figure1-1.png",
        "caption": "Fig. 1. Structure of the LSRA. (a) Cross section. (b) 3-D structure.",
        "texts": [
            " The inductance model can be obtained from finite element analysis (FEA) and polished by experimental measurement, and then the actual model can be incorporated into the control system in forms of look-up table and analytical expressions [2], [10]. To achieve a precise model, look-up table is used in this paper. To avoid the disturbance of measurement noise and error, a novel tracking differentiator (TD) is used to calculate the inductance derivative based on the preobtained phase inductance curves. Nonlinear TD is effective to acquire differential signal in practical system based on optimal solution of the second-order equation [3]. In this paper, the developed LSRA is a four-phase cyclical structure shown in Fig. 1. Each phase of the LSRA is composed of six coils that are placed in a cyclical container, as shown in Fig. 1(a). The compact configuration can generate high force in a limited space that is especially suitable for volume constraint application such as suspension system. To smooth the output force profile, four-phase configuration is selected, as shown in Fig. 1(b). In this paper, model of the LSRA and the inductance profile and force dynamic are given. Nonlinear TD presented in Section III is used to estimate the inductance and its derivative for the purpose of force control. A simple Proportional-Integral (PI) control with compensation unit based on the knowledge of inductance derivative is proposed. The simulation results verify the effectiveness of the force control scheme. Manuscript received March 8, 2014; revised May 23, 2014 and June 23, 2014; accepted June 24, 2014"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001174_amm.465-466.720-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001174_amm.465-466.720-Figure1-1.png",
        "caption": "Fig. 1: Finite element model for laser heating thermal analysis",
        "texts": [
            " For metallurgical analysis, the work-pieces were sectioned and polished perpendicular to irradiation travel direction at 15 mm from the irradiation stating edge. The melted zones (MZ) and heat affected zones (HAZ) were observed using high magnification microscope. Table 1 shows the scanning parameters used in experimental as well as in numerical analysis. Thermal properties of titanium alloy were referred from [4]. Since the scanning process only generates heat on the top of the surface, therefore the keyhole effect can be neglected. Numerical analysis. Fig. 1 shows the schematic diagram of the developed numerical analysis model in this study. The model of laser irradiation was made with circular heat source moves along a 10 mm length of center line. The value of absorption rate A was considered as constant to simplify the model. To reduce the element number, the model was made with half size as the heat source and the actual specimens are symmetric. The value of A was varied between 40 to 60 % by referring to [5]. Total size of the model is 6 mm x 10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002404_mmar.2016.7575286-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002404_mmar.2016.7575286-Figure1-1.png",
        "caption": "Fig. 1. Cameras frame xc1yc1zc1, xc2yc2zc2 and reference frame xyz. The points Pccor1, Pccor2 are corrected images of the corrected point Pcor",
        "texts": [
            " The characteristics of the calculation of the position and orientation of an manipulation object using program dist0 contains the third point. The fourth point describes the calculations of the robot desired trajectories. Example calculations of the desired trajectories is provided in section five. The summary of this work is the sixth point. II. POSITION COORDINATES OF THE CHARACTERISTIC To observe the robot Adept Six-300 environment the system of two cameras Edimax IC-7100 P (described in [5]) was used. The frames xc1yc1zc1, xc2yc2zc2 of these cameras together with the reference frame xyz are illustrated in Fig.1. The reference frame xyz is associated with the robot station. The F1 is focal of the Camera1, and F2 - focal of the Camera2. The Pcor is the point calculated from the coordinates of the images Pc1 and Pc2 of the characteristic point P. These images are read properly from the Camera1 and Camera 2. The real images Pc1 and Pc2 are different from the ideal images Pci1 and Pci2. Points Pci1 and Pci2 lie on a straight passing through the point P and respectively by points F1 or F2. These points are illustrated in Fig",
            " Based on the known dimensions of a single pixel coordinates of the points Pc1 and Pc2 are converted to millimeters. Next these coordinates are converted to cameras frame xc1yc1zc1, or xc2yc2zc2. For these coordinates the errors caused by optical distortion are calculated. For these calculations of the errors an original method of interpolation [5] was used. These errors are included in the calculation of the corrected coordinates cxccor1, cyccor1 of the point Pccor1, in the frame xc1yc1zc1, and the coordinates cxccor2, cyccor2 of the point Pccor2, in the frame xc2yc2zc2 (see Fig. 1). For those coordinates of points Pccor1 and Pccor2 the coordinates xcor, ycor of the point Pcor are calculated in reference frame xyz. For these calculations are needed following coordinates: xccor1, yccor1, zccor1 of the point Pccor1; xccor2, yccor2, zccor2 of the point Pccor2; xF1, yF1, zF1 of the point F1 and xF2, yF2, zF2 of the point F2, in the reference frame xyz. Equations (1a, b) describe these coordinates.               =             1 0 1 ccori c ccori c ci ccori ccori ccori y x z y x T ; i=1,2",
            " The point F1 is on the axis zc1 and is distant from the plane xc1yc1 about focal length fc1. Similarly the point F2 is on the axis zc2 and is distant from the plane xc2yc2 about focal length fc2. Therefore in (1b) the coordinates cxcFi=cycFi=0 and czcFi=fci. Homogeneous transformations matrices Tc1 and Tc2 are calculated after turning on the server. Matrix Tc1 is a homogeneous form of a description of the frame xc1yc1zc1 in reference frame xyz. Similar the matrix Tc2 is a homogeneous form of a description of the frame xc2yc2zc2 in reference frame xyz. Figure 1 shows that the point Pcor lies at the intersection of straight passing through the points Pcor, F1, Pccor1 and straight passing through the points Pcor, F2, Pcor2. These lines describe respectively (2a) and (2b). = \u2212 \u2212 11 1 ccorF ccorcor xx xx = \u2212 \u2212 11 1 ccorF ccorcor yy yy 11 1 ccorF ccorcor zz zz \u2212 \u2212 , (2a) = \u2212 \u2212 22 2 ccorF ccorcor xx xx = \u2212 \u2212 22 2 ccorF ccorcor yy yy 22 2 ccorF ccorcor zz zz \u2212 \u2212 . (2b) These equations allow the creation systems of 3 equations with 3 unknowns xcor, ycor and zcor",
            "9617), P4(468.80 79, 603.2844). For these coordinates were calculated coordinates cxc and cyc of the characteristic point images in the respective cameras frame xc1yc1zc1 or xc2yc2zc2. For these coordinates the program Tobject calculated the optical distortion errors and corrected coordinates cxccor and cyccor. The results of these calculations are presented in Table 1 and 2. The coordinates cxc1 and cyc1 describe the position of the characteristic point images Pc1 in the Camera1 frame xc1yc1zc1 (see Fig.1). Similarly the coordinates cxc2 and cyc2 describe the position of the characteristic point images Pc2 in the Camera2 frame xc2yc2zc2. The coordinates cxccor1 and cyccor1 describe the position of the characteristic point images corrected Pccor1 in the Camera1 frame xc1yc1zc1. Similarly the coordinates cxccor2 and cyccor2 describe the position of the characteristic point images corrected Pccor2 in the Camera2 frame xc2yc2zc2. The focal length of the cameras fc1=fc2=5.02 mm. After turning on the server the matrices Tc1 and Tc2 were calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003083_012017-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003083_012017-Figure2-1.png",
        "caption": "Fig. 2. Final position of the trailer-truck, not equipped with a stabilization system, after emergency braking in a turn",
        "texts": [
            " The simulated maneuver consisted of the following parts: entrance to the turn, the movement in the turn with a constant angle of the steered wheels and an increasing heading speed not exceeding critical speed according to the rollover condition, emergency braking. At the same time, the braking torque of 0.7 from the maximum value was realized at the wheels of the left side of the semi-trailer, and 0.1 of the maximum value of the braking moment at the wheels of the right side. Research of the movement of trailer-truck not equipped with a stabilization system In Figure 2 shows the final position of the trailer-truck, not equipped with a stabilization system, after stopping. Fig. 3 shows the time-history of the hitch angle \u03b3. Design Technologies for Wheeled and Tracked Vehicles (MMBC) 2019 IOP Conf. Series: Materials Science and Engineering 820 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/820/1/012017 Figure 2 and 3 show that during emergency braking in a turn during an emergency failure of the braking system of a semi-trailer truck, deviations from a given trajectory are significant. The hitch angle reaches 750. Research of the movement of trailer-truck equipped with a stabilization system In Figure 4 shows the final position of the trailer-truck, equipped with a stabilization system, after stopping. Fig. 5 shows the time-history of the hitch angle \u03b3. Fig. 4. Final position of the trailer-truck, equipped with a stabilization system, after emergency braking in a turn From the presented simulation results (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000513_0954407015590703-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000513_0954407015590703-Figure2-1.png",
        "caption": "Figure 2. The hammer force and the mesh of test points.",
        "texts": [
            " Since in the ring model the relative motion of the tread with respect to the fixed rim is considered, tests were carried out by fixing the hub of the rim to a very stiff column (Figure 1). Preliminary tests14 showed that the rim modes do not influence the out-of-plane vibrational properties of a tyre in the low-frequency range (below 300Hz). Tests were carried out with the impulsive method. A lateral impulsive force was generated by means of a hammer for modal analysis at six points of the tyre\u2019s carcass evenly spaced by 60 . Figure 2 shows the typical hammer force and the test points. The response was measured by means of a triaxial accelerometer fixed to the lowest point of the tyre (Figure 3). Since the first modes of the tyres are the most important modes, six measurement points were considered a good compromise between the need to hava a sufficient number of points to identify the modal shapes and the need to develop a quick testing method. For each measurement point the frequency response function (FRF) was calculated as the ratio of the crossspectrum of the input signal force and the output signal (acceleration) and the auto-spectrum of the input signal; see the book by Ewins15 for a detailed description of calculation of the FRFs for modal analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure4.77-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure4.77-1.png",
        "caption": "Figure 4.77 Linear motion; ballscrew drive",
        "texts": [
            " Friction Force (FF): This is the opposing force created by the friction between the load and the load bearing surface. Do not confuse the coefficient of friction (\u03bc) with the screw efficiency (e). Nut Preload (TP): To eliminate backlash, the drive nut, through which the screw rotates, is sometimes preloaded. This preload creates an additional torque load on the motor. Note: When performing screw calculations do not confuse the screw pitch (PS) which has the units of rev/cm with the screw lead (LS) which has the units of cm/rev. (See Figure 4.77). 240 Electromechanical Motion Systems: Design and Simulation System Data: JM = motor inertia (g cm s2) JC = coupling inertia (g cm s2) W = load weight (g) TP = preload torque (g cm) FL = load force (g) FG = gravity force = W (g) FF = friction force = \u03bcW (g) JS = screw inertia (g cm s2) TB = bearing torque (g cm) PS = screw pitch (rev/cm) e = screw efficiency \u03bc = friction coefficient Motion: Position: \u03b8M = (2\u03c0 PS)(S) (rad) Velocity: \u03b8 \u2032 M = (2\u03c0 PS)(S\u2032) (rad s\u22121) Acc/Dec: \u03b8 \u2032\u2032 M = (2\u03c0 PS)(S\u2032\u2032) (rad s\u22122) At Motor: TR = TP + TB + \u00b1FL + FF \u00b1 FG 2\u03c0 PSe (reflected torque) JR = ( W 980"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000793_cdc.2013.6760323-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000793_cdc.2013.6760323-Figure1-1.png",
        "caption": "Fig. 1. The motion of the unicycle restricted to the four components of P* (a) p\ufffd,J (b) p\ufffd,r (c) p*,! (d) P*'\ufffd. 't t,+ t,+ 1., 1.,",
        "texts": [
            " Since P* is a product manifold, it is sufficient to characterize Pi for each unicycle. The zero dynamics algorithm applied to (1) with output Qi(Xi) yields Pi = {Xi E IR4 : Qi(Xi) = (3) Ox,Qi cos (8i) + Oy,Qi sin (8i) = 0, Wi + vi- O} . The set Pi is not connected, it consists of four disjoint components. These four components are distinguished by the fact that Wi + V i- 0 on Pi. Each component of Pi corresponds to a distinct type of motion for unicycle i along its path. The four different types of motion are depicted in Figure 1 for unicycle i. In Figure 1 0-; (.\\) is the tangent vector to Ii at 0-i (.\\) and indicates the parameterization direction. Since Wi + V i- 0 on Pi, unicycle i cannot switch between these four situations. Specifications PF is satisfied by making (3) attractive for each unicycle. If path i is closed then IDl; = JR mod Li where Li > 0 is the length of the curve. If path i is non-closed then ][))i = R We now introduce a projection, similar to that used in [12], in the output space of the unicycle that associates to each point Zi sufficiently close to the path Ii a number in ][))i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000142_ecce.2019.8912707-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000142_ecce.2019.8912707-Figure1-1.png",
        "caption": "Fig. 1. Cross section of 1/6 model of 18 \u2013 12 pole switched reluctance motor.",
        "texts": [
            " However, this method is still limited to low speed operation where the back electromotive force (Vemf) is low. Therefore, the current regulation is easy. In this paper, the improved algorithm of current derivation is proposed to realize the flattening radial force sum at middle speed operation. The current profile is derived with the consideration of the back electromotive force Vemf and the limited input DC voltage of the inverter. Therefore, minimum variation of radial fore sum can be obtained although the current profile is influenced by high Vemf. II. RADIAL FORCE IN SWITCHED RELUCTANCE MOTOR Fig. 1 shows a part of cross section of three-phase switched reluctance motor. The number of stator and rotor poles are 18 and 12, respectively. The stator poles have short pitch concentrated windings assigned as A-phase, B-phase, and C-phase. When the current is excited in the stator pole, instantaneous tangential and radial force are generated. The radial force acting in stator poles in A-phase, B-phase, and Cphase are defined by FrA, FrB, and FrC, respectively. 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 5231 According to [13] \u2013 [26], the acoustic noise of switched reluctance motors can be reduced by minimizing variation of radial force sum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002078_msf.843.106-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002078_msf.843.106-Figure1-1.png",
        "caption": "Fig. 1. The installation diagram for EMP treatment of metal melts: 1 \u2013 an EMP generator; 2 \u2013 wires; 3 \u2013 an asbestos cover; 4 \u2013an emitter; 5 \u2013 a protective quartz tube; 6 \u2013 a shaft resistance furnace",
        "texts": [
            " received the first positive results for casting alloys [2]. Shaburova N.A. systematized the results of pulse treatment of ferrous and non ferrous metals [3,4]. Balakirev V.F. et. al developing a theory of NEMP action on metal melts [5,6]. AL9 aluminium alloy of Al-Si system was selected as a material for studying. The chemical composition of AL9 alloy is: 6.4 wt.% Si, 0.06 wt.% Cu, 0.01 wt.% Mn, 0.27 wt.% Mg, 0.02 % Ti, 0.04 wt.% Zn, 0.27 wt.% Fe, 92.9 wt.% Al. The installation diagram for pulse electromagnetic treatment of the metal melt is shown in Fig. 1. Metal melting was carried out in a shaft resistance furnace. Generator (1) creates unipolar pulses: pulse duration \u2013 1 ns , amplitude \u2013 10 kV, pulse repetition frequency \u2013 1 kHz. One contact of the generator was closed on the crucible, the second one was connected with a copper radiator (4) and immersed in the melt. To protect the radiator from its contact with the metal melt, a quartz tube (5) was used. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002264_9781118758571-Figure1.11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002264_9781118758571-Figure1.11-1.png",
        "caption": "FIGURE 1.11 Circular coaxial transmission line and a cross section of the line.",
        "texts": [
            " Substituting these conditions into equation (1.57) gives us two equations in the two unknown constants. Solving for these constants and putting the results back into equation (1.57) gives us identically equation (1.47). In Chapter 2 we will see that in many practical situations a structure is very long and uniform in one dimension and that therefore a cross section of the structure gives us excellent results for V, E, U, and C, the latter two on a per-unit-length basis. An example of such a structure is the circular coaxial cable shown in Figure 1.11. This is an example of a transmission line; transmission lines will be discussed in Chapter 2. The salient point here is that these are two concentric circular conductors of radii a < b. As long as we are not near the ends of the cylinders, the electrostatic problem is essentially a twodimensional problem. Laplace\u2019s equation in cylindrical coordinates with both circular symmetry and no length (z) dependence is as follows, from equation (1.53): r2V = d dr r dV dr = 0 \u00f01:58\u00de Again, integrating twice and using the boundary conditions V(a) = 0 and V(b) = V0, we get V =V0 ln r=a\u00f0 \u00de ln b=a\u00f0 \u00de \u00f01:59\u00de and evaluating equation (1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001392_s11721-013-0084-9-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001392_s11721-013-0084-9-Figure3-1.png",
        "caption": "Fig. 3 (a) The decision making individual at x moving in direction \u03c6\u2032 has a probability of turning to direction \u03c6 based on its interaction with the neighbor located at s that moves in direction \u03b8 . (b) Visual aid for function g\u03c3a : the decision making individual at x moving in direction \u03c6\u2032 has a probability to turn (due to attraction) to a direction near \u03c6, with maximum turning possibility of ka and with uncertainty \u03c3a",
        "texts": [
            " The integral in the righthand-side is taken over s and \u03b8 , which accounts for all neighbors\u2019 positions and directions, respectively. The other two components of \u03bb are defined similarly: \u03bbr( x,\u03c6) = qr \u222b R2 \u222b \u03c0 \u2212\u03c0 Kd r ( x \u2212 s)Ko r ( s; x,\u03c6)u( s, \u03b8, t) d\u03b8 d s, (10) \u03bbal( x,\u03c6) = qal \u222b R2 \u222b \u03c0 \u2212\u03c0 Kd al( x \u2212 s)Ko al(\u03b8,\u03c6)u( s, \u03b8, t) d\u03b8 d s, (11) where qr and qal are strengths of repulsion and alignment, respectively. Reorientation term T ( x,\u03c6\u2032, \u03c6) The reorientation term T ( x,\u03c6\u2032, \u03c6) describes the rate at which an individual located at x reorients itself from \u03c6\u2032 to \u03c6 due to influences of its neighbors\u2014see Fig. 3(a). Similar to \u03bb, this term also takes contributions from attraction, repulsion and alignment interactions: T ( x,\u03c6\u2032, \u03c6 ) = Ta ( x,\u03c6\u2032, \u03c6 ) + Tr ( x,\u03c6\u2032, \u03c6 ) + Tal ( x,\u03c6\u2032, \u03c6 ) . (12) The contribution from attraction is taken to be: Ta ( x,\u03c6\u2032, \u03c6 ) = qa \u222b R2 \u222b \u03c0 \u2212\u03c0 Kd a ( x \u2212 s)Ko a ( s; x,\u03c6\u2032)wa ( \u03c6\u2032 \u2212 \u03c6,\u03c6\u2032 \u2212 \u03c8 ) u( s, \u03b8, t) d\u03b8 d s, (13) where qa and the interaction kernels are the same as before, and wa is a turning probability function that gives the probability of turning from \u03c6\u2032 to \u03c6 due to attractive interactions with the neighbor located at s",
            " (35) The probability function wa is modeled by wa ( \u03c6\u2032 \u2212 \u03c6,\u03c6\u2032 \u2212 \u03c8 ) = g\u03c3a ( \u03c6\u2032 \u2212 \u03c6 \u2212 va ( \u03c6\u2032 \u2212 \u03c8 )) , (36) where g\u03c3a is an approximation of the delta function with width \u03c3a , and va is a turning function. The decision making individual can turn to any direction within a specific range. This range is centered around the direction \u03c6 = \u03c6\u2032 \u2212 va ( \u03c6\u2032 \u2212 \u03c8 ) , which happens when the argument of the function g\u03c3a is zero. The parameter \u03c3a > 0 measures the width of the turning range the decision making individual will move into\u2014see Fig. 3(b). The smaller the \u03c3a , the more accurate the turning. If \u03c3a is large, then the range is wide and the decision making individual can move anywhere within the range. The function to describe g\u03c3a is taken to be g\u03c3a (\u03b7) = 1\u221a \u03c0\u03c3a \u2211 z\u2208Z e \u2212( \u03b7+2\u03c0z \u03c3a )2 , (37) which is a periodic Gaussian with extra contributions from full rotations. The turning function va is modeled as va(\u03b7) = ka sin\u03b7, (38) where ka is a constant between 0 and 1 that describes how much the decision making individual will turn due to attraction (see Fig. 3(b)). An alternative choice is va(\u03b7) = ka\u03b7, which may be more biologically realistic. However the choice (38) for va is made because it is periodic and works well with the fast Fourier transform, which the numerics in this paper is based on, as discussed in Sect. 3. The probability functions wr and wal are defined through the same steps as equations (36)\u2013(38), using approximations g\u03c3r , g\u03c3al to the delta function, with widths \u03c3r , \u03c3al, turning functions vr , val, and turning strengths kr , kal, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003522_s00542-020-05045-8-Figure8-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003522_s00542-020-05045-8-Figure8-1.png",
        "caption": "Fig. 8 Few example of buckling modes with fingertip is freely moving (no grasping). a First buckling mode in open mode under EM actuation. b First buckling modes in close mode under EM actuation. c Second buckling mode in close mode under PE actuation",
        "texts": [
            " In this section, the buckling load and buckling modes are evaluated for freely moving fingertips (no grasped object), fixed end (grasping object with no slippage), and frictionless support (grasping object with no friction). The first buckling scenario is studied for open mode under EM input force. The simulation is carried out for Input load of 0.04 N applied in the upward direction of shuttle with the anchors and the package base are fixed. Upon simulation, the load multipliers for the first two modes of buckling are obtained in Fig. 8. The buckling load Pb is equal to M x P. Where M is the load multiplier, and P is the applied input load. The two buckling modes took place on the beam (L5) that connects package to the central shuttle. An example of the first buckling mode shape is illustrated in Fig. 8a. Here, the deformation values have no significance, since they are scaled by ANSYS Workbench. Only the mode shapes are considered useful. The second buckling scenario assumes fingertips movement in close mode are free and the gripper is actuated by EM input force. The input load of 0.002784 N (corresponds to displacement of 0.208811 mm which is the maximum safe input load just before collision of fingertips) is taken, and then is applied in the downward direction onto (a) (b) (c) (d) S = 0.0073F - 1E-060 0",
            "04 0 100 200 D is pl ac em en t (m m ) Force (N) Fig. 7 Close mode under Piezo-electric force: a Equivalent stress for input force of 190 N. b Plot of equivalent stress and force. c Directional deformation for input force of 190 N. d Plot of directional deformation and force the center of shuttle with anchors and package base being fixed. The output is obtained in the form of load multipliers and results are outlined in Table 5. The first three buckling modes took place on the long beams (L1) with the example in Fig. 8b, while the last three modes are due to buckling in the connecting beams (L5). The third buckling scenario is studied for free fingertips moving in the closing mode with gripper being actuated by PE input force. Upward input load of 190 N is applied on the anchor with both central shuttle and the package base are fixed. The output load multipliers are obtained and the buckling loads are calculated in Table 6. The first three modes occurred at (L1) beams with example shown in Fig. 8c, and the remaining modes occurred along the wings which connects the anchor with the package base. Another group of buckling scenarios studies conditions when the gripper is holding an object. In this example the fingertips are fixed in the FEM model, and the loading and constraint conditions used in the last three scenarios are reapplied. In the close mode under EM actuation, the first three buckling modes in Table 6 occurred at the end of (L3) beams with a sample simulation shown in Fig. 9a. The simulation results of the close mode under the PE actuation are summarized in Table 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003117_s00542-020-04893-8-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003117_s00542-020-04893-8-Figure11-1.png",
        "caption": "Fig. 11 Experimental setup for testing the VSY-SEA performances",
        "texts": [
            " 7 is 1, 2, and 4, the resulting stiffness curve is linear, quadratic, and quaternary polynomial. For the experiments, a VSY-SEA prototype was made, as shown in Fig. 10. This model is a prototype of the VSYSEA whose purpose is verifying the design, so the major material of the prototype is just aluminum. Later, for further study, the model will be made of a material that minimizes wear by rollers. As mentioned earlier, the stiffness of the linear spring used in VSY-SEA is 5.9 N/ mm, and a load arm is installed for measuring an external torque. As shown in Fig. 11, the external force was applied and measured by the linear stage and force transducer, and the rotation angle was measured using the rotary encoder. As in the simulation, the experiments were carried out in two cases when the value of n was fixed, and when the value of kt was fixed. The experimental results are shown in Figs. 12 and 13. In the first case, as shown in the simulation, the stiffness curves of the system are linear for all kt values, as shown in Fig. 12. The stiffness of the system was obtained using linear interpolation, and the stiffness of VSY-SEA increased from 341"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002027_j.protcy.2016.03.005-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002027_j.protcy.2016.03.005-Figure1-1.png",
        "caption": "Fig 1. Working principle of the Radial Magnetic Bearing",
        "texts": [
            " The review of the past literature ravealed that considerable amount of work is still required to study the control aspects of four pole pair radial magnetic bearing. Hence, in this work it is attempted to look for modelling, formulation and system response which are shown in section (2, 3). Further an analysis of four pole pair radial magnetic bearing was performed considering PID controller, simulink results and conclusions are detailed in sections (4, 5). The basic layout of a magnetic bearing system is shown in Figure1. Stationary electromagnets are positioned around the rotating assembly of a machine. Typically, two radial magnetic bearings are used to support and position the shaft in the lateral (radial) directions and one thrust bearing is used to support and position the shaft along the longitudinal (axial) direction. When the magnetic bearing is operating, each magnetic bearing rotor is ideally centred in the corresponding stator so that contact does not occur. The position of the shaft is controlled using a closed-loop feedback system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003248_j.cja.2020.06.030-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003248_j.cja.2020.06.030-Figure1-1.png",
        "caption": "Fig. 1 Radical relief grinding movement of a grinding wheel.",
        "texts": [
            " Discussed are conditions of: cutting-tool angle, including zero or positive rake angle; hand of helix, including right-handed, lefthanded or annular buckle which is non-handed. Finally, clearly identify the validity of the developed method by using the numerical and experimental examples of the gear hob grinding. In order to analyze the radical relief grinding movement of grinding wheels correspond to gear hobs, the model based on the spatial motion relationship of the above two objects in an arbitrary moment of the machining process is constructed in Fig. 1. where SH XH;YH;ZH\u00f0 \u00de and SW XW;YW;ZW\u00f0 \u00de are respectively the moving coordinate systems fixed to the gear hob and grinding wheel. ZH and ZW are respectively the axes coincides with hob shaft and grinding wheel shaft. OH and OW are the axis midpoints of the gear hob and grinding wheel. The key to calculating the grinding wheel axial profile is the analysis on relative location of two objects. As shown in Fig. 2, the workpiece that needs to be ground is a right-handed gear hob and in Fig. 3, it is an annular buckle gear hob"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002160_icra.2016.7487445-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002160_icra.2016.7487445-Figure4-1.png",
        "caption": "Fig. 4. Measurement device for stress distribution measurements of a wheel with grousers.",
        "texts": [
            " The friction between the guide plate and the top portion of the guide pole was minimized by using ball-casters, which were fixed to the pole at the contact point with the guide plate. The pulleys and the rope were installed to the sandbox, and the wheel slippage of the testbed is induced by adding traction load to the testbed via the rope. Furthermore, we used a motion capture system (Osprey, Motion Analysis Corporation) for tracking the rover motion. The traveling velocity, the wheel sinkage, and the wheel rotation speed were obtained from motion capture system by attaching markers to the testbed. As shown in Fig. 4, we measured the three-dimensional forces acting on the wheel with grousers using four configurations. One was the configuration for measuring the force acting on the grouser (Fig. 4 (a)), and the others were the configurations for measuring the force acting on the wheel surface between each grousers (Fig. 4 (b), (c), (d)). The measurement locations at the middle of wheel surface and at the grouser have a phase difference of 15\u25e6 from each other. That is, the measurement location for the middle of wheel surface (Configuration B) is exactly between two grousers. The measurement locations at the wheel surface near the grouser (Configuration C and D) and at the grouser have a phase difference of 5\u25e6 from each other. Here, contactparts to measure the forces on all the wheel surfaces and the grouser can be treated as rigid parts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001788_icems.2015.7385248-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001788_icems.2015.7385248-Figure1-1.png",
        "caption": "Fig. 1. Principle sketch of the stator cage machine.",
        "texts": [
            "78-1-4799-8805-1/15/$31.00 \u00a92015 IEEE I. INTRODUCTION Recently, a new kind of electrical machine, so-called \u201cStator Cage Machine\u201d, has been developed [1]. The stator consists of a stack of iron lamination (like for conventional electrical machines), but with massive conductors in each slot, being short-circuited at one axial end of the machine, see Fig. 1. Each slot conductor is connected to the center tap of a half bridge, so that the current in each slot can be determined individually. The only boundary condition for the stator slot currents is that the sum of all stator slot currents is equal to zero. Using a squirrel-cage rotor, it is possible to change the number of pole pairs during operation of the machine. Therefore, in the following an analytical calculation method will be developed considering this flexibility. II. MAGNETO-MOTIVE FORCE OF THE STATOR Usually the stator magneto-motive force (MMF) is calculated by summing up the MMF distribution of every single turn, see e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003243_tia.2020.3008369-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003243_tia.2020.3008369-Figure4-1.png",
        "caption": "Fig. 4. Magnetic circuit distribution model of SPM As shown in Fig. 4, ra is the mid-diameter radius of air-gap, ht is the tooth height, rs() is the radius of the magnetic circuit curve at the stator yoke, rr() is the curve of the magnetic circuit curve at the rotor yoke, wt is the stator tooth width, hs is the height of stator yoke.",
        "texts": [
            " l() is expressed as: Authorized licensed use limited to: University of Exeter. Downloaded on July 15,2020 at 05:42:49 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (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. IEEE TRANSECTION ON INDUSTRY APPLICATION 5 The idea of distributed magnetic circuit method can be used for the calculation of saturation coefficient. As shown in Fig. 4, the motor can be radially divided into five regions: air-gap region (I), stator tooth region (II), stator yoke region (III), PM pole region (IV), and rotor yoke region (V). In the circumferential direction of the magnetic circuit, the spatial mechanical angle  in Fig. 4 can be used as a symbol for the position. According to Fig. 4, the motor magnetic circuit satisfies the following: ( ) ( ) ( ) ( ) ( )m g t s rF F F F F        (35) where Fm() is the magnetomotive force generated by region IV, Fg() is the air-gap magnetic voltage drop of region I, Ft() is the stator tooth magnetic voltage drop of region II, Fs() is the stator yoke magnetic voltage drop of region III, and Fr() is the rotor yoke magnetic voltage drop of region V. To obtain the magnetic voltage drop of each part of the motor, the magnetic density of each part needs to be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001040_j.mporth.2016.05.014-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001040_j.mporth.2016.05.014-Figure1-1.png",
        "caption": "Figure 1 (a) Moment of force F about point X: magnitude\u00bcFa, direction\u00bcout of page. (b) Moment of force W about point X: magnitude\u00bcFbcos q, direction\u00bcinto page. (c) Force of mass due to gravity\u00bcmg, perpendicular distance from mass to shoulder\u00bcd sinf, magnitude of moment on shoulder joint due to mass\u00bcmgd sinf.",
        "texts": [
            " The SI unit of force is the Newton (N) which is equivalent to 1 kgms 2. Since force is a vector, both the magnitude of the force and the direction in which it acts are important. The moment of a force (M) is defined as the product of the force and the perpendicular distance about which it rotates. Moment is also a vector quantity; the direction of a moment is described by the \u2018right-hand rule\u2019: if the fingers of the hand are curled to represent the direction of rotation then the thumb, if held perpendicular to the fingers, represents the direction of the moment (Fig. 1). The unit of moment is the Newton-metre (Nm). Often in mechanics, we wish to add physical quantities together. With scalars, this is done simply by summing the magnitudes, but with vectors, the direction must also be taken into account. With two or more vectors in two dimensions, this can be done most easily using the parallelogram rule. Here, each vector quantity is depicted by an arrow in the appropriate direction, the length of which is proportional to the magnitude of the vector. The arrows form two sides of a parallelogram and the sum or resultant vector is the diagonal of the parallelogram (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.30-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.30-1.png",
        "caption": "Figure 3.30 Ironless linear motor schematic",
        "texts": [
            " The air gap between the forcer and the magnet assembly is critical in establishing the proper level of the magnet field strength. The bearing structure and the flatness of the surface on which the magnet assembly is mounted must be tightly controlled. For example, a variation of 30% in the air gap can produce a 5% variation in the force constant that can create a low frequency force constant ripple determined by the pitch of the flatness variation. Bymounting two sets of linearmagnet assemblies opposite each other, as shown in Figure 3.30, a \u201cU\u201d-shaped structure is created. The forcer is located in the air gap between the two magnet assemblies. 78 Electromechanical Motion Systems: Design and Simulation The forcer, consisting of a molded three phase coil set, contains no iron, resulting in zero cogging. The forcer design is analogous to the printed circuit disc and basket weave armatures in low inertia rotary motors. In addition, there is no magnetic attraction force as in the iron motor, reducing bearing loading. The result is a forcer with lower mass than a comparable iron based forcer, providing high acceleration capability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.75-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.75-1.png",
        "caption": "Figure 3.75 Cycle rate versus single cycle energy",
        "texts": [
            " If operated at the specified voltage, the device will remain well within its temperature rating. In high cycling rates the load energy being transmitted and absorbed will cause the temperature to rise. Unlike the case for motors, where a published thermal resistance together with a calculated RMS dissipation allows the internal temperature to be determined, the nonlinear nature of a clutch or brake does not permit this. Instead, manufacturers provide charts showing the maximum permissible cycle rates in two ways: (a) Energy per cycle versus cycles per minute (Figure 3.75) (b) Inertia versus cycles per minute for various speeds (Figure 3.76). For (a) the energy per cycle, K, must first be calculated by: K = J\u03b8 \u20322 2 g cm (3.119) then the cycle rate determined from the chart. Note that manufacturers use different units for this expression. Be sure to convert your units to those used on the chart. For (b), use the total inertia and the maximum operating speed to determine the maximum cycling rate. Both methods will provide the maximum allowable cycling rate that will result in the devices not exceeding their maximum rated temperature in a 25 \u25e6C ambient"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002371_978-981-10-2309-5-Figure1.1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002371_978-981-10-2309-5-Figure1.1-1.png",
        "caption": "Fig. 1.1 Evolution of rotary motor to linear machine",
        "texts": [
            " 123 Contents xv Acronyms 2/3D Two/three-dimensional AC Alternating current BCs Boundary conditions CM Coenergy method CNC Computer numerical control DC Direct current EALS Electromagnetic aircraft launch system EMC Equivalent magnetic circuit FE Finite element FEM Finite element method LFE Longitude fringe effect MMF Magnetomotive force MST Maxwell stress tensor PC Personal computer PM Permanent magnet PMLMs Permanent magnet linear machines PMTLM Permanent magnet tubular linear machine RMS Root mean square xvii List of Figures Figure 1.1 Evolution of rotary motor to linear machine . . . . . . . . . . . . 2 Figure 1.2 Linear machine for electromagnetic catapult [8] . . . . . . . . . . 3 Figure 1.3 Linear machines for applications of logistics [11]. . . . . . . . . 4 Figure 1.4 Linear machines for applications of Industrial Automation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 Figure 1.5 Linear machines for applications of rail transportation . . . . . 6 Figure 1.6 Winding arrangements for tubular PM linear machines [19] ",
            " After 1970s, SIEMENS Company, Westinghouse Company and Panasonic Company have paid more attention on the investment of linear machines, and a large number of linear machines with different features have emerged in market. \u00a9 Springer Science+Business Media Singapore 2017 L. Yan et al., Electromagnetic Linear Machines with Dual Halbach Array, DOI 10.1007/978-981-10-2309-5_1 1 2 1 Introduction An electric machine is a device that converts electricity into mechanical motions. The linear machine can be considered as one evolution of rotary machines in the structure as shown in Fig. 1.1. Permanent magnets (PMs) are mounted on the stator and windings are on the mover, or vice versa. The interaction between magnetic field of PM poles and current input in the windings generates force in the unrolled direction. Compared with conventional linear drive system consisting of rotary motors and ball screws, linear machines have many merits as follows: \u2022 Linear machines are direct drive systems that do not need ball screws, cranks, gears, or any other motion conversion mechanisms, which helps to achieve compact size of system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000104_s12206-019-1038-y-Figure23-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000104_s12206-019-1038-y-Figure23-1.png",
        "caption": "Fig. 23. Hydraulic TPM with the workpiece (1. Base platform, 2. Moving platform, 3. Hydraulic cylinder, 4. Position sensor, 5. Control valves, 6. Rotary connections, 7. Planar joint, 8. Forging).",
        "texts": [
            " The proportional gain Kp and the period Tp are calculated by equations [28]: 1 2 3 1 2 3 1 2 p p a a aK b b b T p w - + -\u00ec =\u00ef - +\u00ef \u00ed \u00ef =\u00ef\u00ee (22) where a1, a2,, a3, b1, b2, b3 are estimated parameters. Fig. 22 shows the graph plots of poles locations for stable position trajectories of the IEHSD. The introduced modified controller tuning method avoids instability of the control system in the case of critical parameters. To assess the effectiveness of synchronous control of three IEHSDs, an experiment was carried out with a TPM loaded with the workpiece (forging), as shown in Fig. 23. Positioning errors and synchronization errors of each IEHSDs affect the tracking accuracy of the TPM. A block diagram of TPM control with three IEHSDs is shown in Fig. 24. The IEHSDs control systems have an internal closed-loop position control system and an external closed-loop synchronization control system. The synchronization control strategy aims to make the differential position errors of the three IEHSDs converge to zero. Then the position errors of the i-th IEHSDs are defined as follows: ( ) ( ) ( )d i i mite L t L t= - , (23) where ( )d iL t is the desired generated length and Lim(t) is the actual measured displacements of i-th"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003753_012060-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003753_012060-Figure17-1.png",
        "caption": "Figure 17. Velocity field contours for drum diameter 650 mm.",
        "texts": [],
        "surrounding_texts": [
            "Table 1 and the graph (Figure 14) show the calculations of the maximum speed of the velocity fields at fixed drum and beater revolutions 640 rpm and 2100 rpm, respectively, and different values of the drum diameter. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 At drum diameters less than 550 mm, the maximum speed values are preserved, since the maximum flow rates in these cases are already set by the beater and remain unchanged. Figure 14. Velocity field contours for drum diameter 350 mm. Figure 15. Velocity field contours for a drum diameter of 450 mm. ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060 ERSME 2020 IOP Conf. Series: Materials Science and Engineering 1001 (2020) 012060 IOP Publishing doi:10.1088/1757-899X/1001/1/012060"
        ]
    },
    {
        "image_filename": "designv11_30_0000682_demped.2015.7303706-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000682_demped.2015.7303706-Figure2-1.png",
        "caption": "Fig. 2 depicts a ball bearing where c is the clearance between shaft and inner ring. In addition, the shaft, inner ring and cage speeds are denoted as nshaft, nir and nc respectively.",
        "texts": [],
        "surrounding_texts": [
            "Vibration analysis has been the most wide-spread condition monitoring technique applied on wind turbines generators employing accelerometers installed radially at the load zone, as shown in Fig. 1 [10]. Tracking of speed related spectral components describing the shaft dynamics, such as the first and higher orders running speeds, is usually implemented in condition monitoring systems along with broadband measurements in various frequency ranges for overall vibration evaluation and early stage bearing defects [7]. Alternatively, an envelope can be set over the considered healthy vibration signature, where an alarm is triggered when a frequency component exceeds the above mentioned predefined limit. In addition to vibration based monitoring, temperature sensors are placed on the bearings\u2019 housings in order to assess the condition of the bearings from a thermal standpoint. Due to the location of the temperature sensors, the measured values reflect approximately the outer ring temperature, when in fact a temperature difference in the range of 10\u25e6C up to 40\u25e6C is expected between the outer and inner rings. Other techniques applicable on generator condition monitoring, such as current signature analysis [11] - [14] or utilization of thermal imaging cameras [15], have been proposed, but with limited field applications."
        ]
    },
    {
        "image_filename": "designv11_30_0000380_iros.2011.6095187-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000380_iros.2011.6095187-Figure3-1.png",
        "caption": "Fig. 3. Trajectory tracking control errors",
        "texts": [
            " Using differentiation of (1), the dynamics of the wall-climbing robot is described as follows, Mq\u0308+D(q\u0307,\u03b8) = MJ\u22121 p\u0308+D(q\u0307,\u03b8) = \u03c4 , (2) where \u03c4 = [\u03c4L,\u03c4R] T and \u03c4L, \u03c4R are wheels\u2019 driving torques. M is the inertia matrix and D(q\u0307,\u03b8) is the nonlinear term which means influences of gravity and other dynamical effects [5]. Equation (2) expresses an inverse dynamics of the wallclimbing robot. This subsection proposes a trajectory tracking control method using the dynamics equation presented in the previous subsection. In Fig.3, W x, W y, \u03b8 and xd , yd , \u03b8d represent coordinates of the robot and their desired ones in \u03a3W . Desired velocity Vd and angular velocity \u2126d are obtained by differentiating xd , yd , \u03b8d . Xe, Ye, \u03b8e represent tracking errors between real position/orientation of the robot and desired ones. Then \u2212\u03c0 < \u03b8e \u2264 \u03c0 . Some trajectory tracking control methods for wheeled mobile robot [6][7] have been proposed. Most of them are kinematics level and assume that the robot moves on horizontal flat surface. They also assume perfect velocity control of the robot in the proof of the convergence of the tracking errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002618_ecc.2016.7810466-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002618_ecc.2016.7810466-Figure1-1.png",
        "caption": "Fig. 1. Vessel earth-fixed and body-fixed frames.",
        "texts": [],
        "surrounding_texts": [
            "The vessel is fitted with two screw propellers and two rudders. In navigation mode, propulsion signals are coupled and, for such a reason, they are supposed to generate identical thrusts. The interaction between rudder and screw propeller is modeled as the linear superposition of the screw propeller and the rudder action. Then, the total delivered force and moment are expressed as the sum \u03c4D = \u03c4P + \u03c4R . (7) In particular, the action of each rudder is represented by a corresponding force and moment array as follows: \u03c4R(\u03b4D, \u03bd) = [\u03c4R,x(\u03b4D, \u03bd), \u03c4R,y(\u03b4D, \u03bd), \u03c4R,n(\u03b4D, \u03bd)]> , (8) which depends on the delivered rudder angle \u03b4D and the velocity array \u03bd. In more detail, each component of \u03c4r is modeled by the lifting-line theory, as follows: \u03c4R,x = \u2212\u03c1CD(|\u03b4D|)AR U2/2 \u03c4R,y = \u03c1CL(|\u03b4D|)AR sgn(\u03b4D)U2/2 \u03c4R,n = YR xR where \u03c1 is the water density, AR is the rudder area, U =\u221a u2 + v2 is the vessel speed, and xR is the longitudinal coordinate of the rudder center of pressure. The delivered thrust by each propeller is computed through the formula: \u03c4P = [KT (u, nD, \u03d5) \u03c1n2DD 2, 0, 0]> (9) where nD is the delivered shaft speed, D is the propeller diameter, KT is a thrust coefficient depending on surge speed, rotary shaft speed, and pitch angle \u03d5. Due to the mechanical symmetries of the vessel, and to the fact that the propellers are running in opposite directions, the lateral forces (second component of \u03c4P ) and the corresponding moments (third component of \u03c4P ) are zero in (9)."
        ]
    },
    {
        "image_filename": "designv11_30_0002989_s12541-020-00344-6-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002989_s12541-020-00344-6-Figure4-1.png",
        "caption": "Fig. 4 a Carved turn mechanism. b When the direction of the skiing coincides with the direction of the slope. c When the direction of the skiing is orthogonal with the direction of the slope",
        "texts": [
            " The ski plate bends in the form of an arc when the edge of the ski plate touches the slope and the turn performs in accordance with the shape of this arc. At this time, the turning radius r is expressed by Eq. (1)\u00a0[18]. Figure\u00a03 shows the parameters of the ski plate and angle . Table\u00a03 is a parameter related to the ski robot\u2019s carved turn mechanism. According to Eq. (1), if we know the ski plate\u2019s L,\u00a0h and controllable , we can control the turning radius. There are two major ways that humans make the carved turn. As shown in Fig.\u00a0 4a, the first way is the abduction\u2013adduction method using the hip roll joint, and the second is the extension\u2013flexion method using the hip pitch, knee, and ankle pitch joints. The abduction\u2013adduction method easily makes the , but it has the disadvantage of over-torque in the hip roll joints. The extension\u2013flexion method also easily makes the , but each joint requires less torque than that of the abduction\u2013adduction method because it uses multiple joints. Previous studies of humanoid robots using the carved turn demonstrated that it is efficient to use both methods [8]",
            " We assumed that the carved turn performed by the robot has no slip phenomenon, the radius of turning is 7 m. At this point, the model of the robot is an inverted pendulum model, and the force acting on the robot is gravity and centrifugal force. It is assumed that the robot\u2019s skiing speed is 5 m/s because the average skiing speed of RoK-2 is experimentally confirmed to be 3 m/s \u223c 5 m/s. As shown in Table\u00a05, we confirmed that the average speed was also about 5 m/s when the robot passes the course in the actual competition. Figure\u00a04b, c show that the robot performs the carved turn on the slope, at which point, the equation of motion can be assumed as Eq. (5). The w is the angle that the robot tilted from a vertical line on the ground when it performs the carved turn. Based on Eq. (5) and the various assumptions, it have be confirmed that the ZMPY of the robot is in the support polygon. The type of the ski robot\u2019s carved turn can be defined by the direction of skiing. One type is when the direction of the skiing coincides with the direction of the slope, and the other is when the direction of the skiing is orthogonal ( 90\u25e6 ) with the direction of the slope. The reason for categorizing according to the skiing direction is that the w is related to the slope angle when the skiing direction is turned 90\u25e6 as in Fig.\u00a04c. When the direction of the skiing coincides with the direction of the slope, it is assumed that the robot is in the right turn as shown in Fig.\u00a04b. w can be assumed as Eq. (6), and the ZMPY of the robot can be defined as Eq. (7). The mass m of the robot is 26 kg, the distance l from the support point to the center of gravity is 0.70806 m, the skiing speed v of the robot is 5 m/s, the radius r of the carved turn performed by the robot is 7 m, gravity acceleration g is 9.81 m\u2215s2 . At this time, the ZMPY of the robot is -0.091 m. The area of the support polygon is defined as Eq. (8), (5)ml2\ud835\udf03w = mglsin\ud835\udf03w \u2212 ( m v2 r ) lcos\ud835\udf03w as shown in Fig.\u00a04a. The distance d between the two feet of the robot is 220 mm. When the ZMP (Eq. (7)) of the robot is compared with the support polygon (Eq. (8)) of the robot, it is confirmed that the ZMPY of the robot is between both feet. Therefore, the robot does not easily lose balance. As shown in Fig.\u00a04c, when the robot is right turning and the direction of skiing is turned 90\u25e6 , w is defined as Eq. (9) due to . At this time, the ZMPY of the robot on the slope can be expressed as Eq. (10). When the ZMP (Eq. (10)) of the robot is compared with the support polygon (Eq. (8)) of the robot, it is confirmed that the ZMPY of the robot is in the support polygon of the robot. In both cases, the stability of the ski turning motion is verified by described assumption and formula. In addition, because the actual skiing speed of the robot is not overly increased because of the frictional force of the snow, the risk of the robot falling over due to centrifugal force is reduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002926_s1068798x20010141-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002926_s1068798x20010141-Figure2-1.png",
        "caption": "Fig. 2. Prototype of the proposed probe.",
        "texts": [
            " With action at some angle \u03b1, the radius of tip introduction is determined from the formula a = R/cos\u03b1. Accordingly, the area of waveguide-tip introduction is determined from the formula S1 = \u03c0a2. The relative error in determining the contact area may be calculated from the formula \u03b4S = tan2\u03b1 \u00d7 100%. Assuming that is approximately 10\u00b0, we obtain . Determining the relative error \u03b4P in pressure measurement from \u03b4F and \u03b4S when \u03b1 = 10\u00b0, we conclude that it is no more than 4.6%. The main component of the prototype probe (Fig. 2) is a force sensor based on tensoresistive measurements. The sensor consists of an elastic element 1 (manufactured from aluminum alloy), to which four tensoresistors 2 are cemented. The sensor is intended to measure forces in the range 0\u20131 N. Its resolution is 0.001 N. Two screws 3 connect adapter 4 to the force sensor. A special threaded aperture in the adapter accommodates the mount for waveguide 5. The force sensor is attached to the panel 6 by two screws 7. Screws 8 attach the upper panel 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003522_s00542-020-05045-8-Figure17-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003522_s00542-020-05045-8-Figure17-1.png",
        "caption": "Fig. 17 Contact simulation of the gripper with spherical object: a total deformation. b Equivalent Von\u2013Mises stress",
        "texts": [
            " 16b showing that the object adheres to the fingertips at the contact region. The contact status becomes a sliding contact as the relative position between the object and fingertips moves away from this region. The pressure at the contact region is also evaluated with results shown in Fig. 16c, and shows that the maximum compressive and tensile pressures are 89 MPa and 46.2 MP, respectively. The maximum static structural deformation is also simulated for the above contact conditions, where the results in Fig. 17a shows large bending of (L3) beam with maximum total deflection of 0.49 mm. The rigidity of the spherical object caused a maximum stress on the fingertip with value of 35.1 MPa as shown in Fig. 17b. 3.8 Gripper scalability The effect of the gripper size is studied in relationship to the displacement of the fingertips and the maximums stress due to input displacement. The overall dimension of the gripper that was studied throughout the previous sections are 20.553 9 17.799 9 37.557 mm3 The structural analysis in Sect. 4 is repeated for gripper size scaled up by a factor of 1:3, 1:5, 1:7, and 1:10. Two cases are considered: In the first case, the magnitude of input displacement of 3 mm is kept constant for each scaled size"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.20-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002662_b978-0-12-800774-7.00002-7-Figure2.20-1.png",
        "caption": "Figure 2.20 EEG electrodes: (A) lateral view and (C) bottom view of the hollow cup showing the wide edge where the colloidal paste is applied to fix the electrode on the scalp. (B) Top view showing the hole to apply the electrolytic gel.",
        "texts": [
            " Elastic bands or adhesive tape are used to attach the electrodes on the skin. Other types of recording electrodes also use metal discs, like the top-hat electrode, that will be discussed in floating electrodes, or formats similar to a disc like EEG electrodes. EEG electrodes are manufactured as stamped, cast, or sintered electrodes. The stamped EEG electrodes, used in routine examinations, are made from a thin plate of silver, gold, or Ag AgCl. They are also named stamped cup electrodes due to their format. They have a wide edge (Figure 2.20A and C) where a layer of colloidal paste is applied to fix the electrodes in the scalp skin; a hole in the top of the electrode (Figure 2.20B) allows easy filling of the electrolytic gel after adhesion of the cup to scalp (and also any excess gel to get out during the positioning of the electrode); and a lateral connector, usually stamped together with the disc, where the lead wire is welded (Figure 2.20A C). EEG stamped discs are sold as individual electrodes or mounted in colored flat cable, which facilitates their application on the scalp with the 10 20 system. The discs are fabricated with different external diameter (6 and 10 mm), central hole diameter (1.5 and 2 mm), and cable lengths (100, 150, and 200 mm) to suit to adult and infant utilization. Cast disc EEG electrodes made of gold usually are cast of pure silver with heavy gold plating. EEG cups are also manufactured as disposable electrodes and in this case they are usually made of Ag AgCl"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000552_icar.2015.7251503-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000552_icar.2015.7251503-Figure3-1.png",
        "caption": "Fig. 3. 3D deformable object model",
        "texts": [
            " We suppose that the external forces are only due to the gravity acting in the negative Z direction according to the reference frame R f (see Fig. 6). The contact manipulation forces are determined with respect to the relative positions and velocities between the fingertips and the external facets of the object mesh. This approach enables to deal with the geometric complexity of the contact surface without the need of a conformal mesh. Figure. 2 illustrates the problem of grasping deformable objects, broken down into its component steps. In this paper, the object presented in Fig.3 is modeled by a nonlinear mass-spring system based on a tetrahedral mesh of the volume: lumped masses are attached to the nodes and spring-dampers elements represent the edges [17] [18]. In this method, springs can be linear or nonlinear according to material type of the object. In our case the springs are linear, but the global behavior of the deformable object is nonlinear due to large displacements and rotations. Tracking the positions of the mesh nodes enables updating the overall object shape as well as the deformation of the contact areas"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003470_j.jnnfm.2020.104406-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003470_j.jnnfm.2020.104406-Figure14-1.png",
        "caption": "Fig. 14. Schematic drawings of the test sections highlighting the up and downstream location of the pressure taps, equidistant (at a distance of about 0.5 mm) from the centre of the [a] circular and [b] elliptic cylinders. Streak images of viscoelastic flows (El \u223c 19).",
        "texts": [
            "5 mm up and downstream of the respective axial stagnation points of the cylinders (this is more than 10\ud835\udc451 in each direction) somehow lead to the measurements being influenced by those effects (since these are typically felt up to several radii downstream)? The distance of about 0.5 mm is believed to be both long enough to mitigate any major influence of such effects on the readings, and close enough to capture the signal of the pressure gradient reasonably well without it being lost as the flow progresses (this would be the case if the pressure taps had been placed far from the obstacles e.g. at the inlet and outlet of the microchannels). Fig. 14 shows a top view of the test sections to give a sense of the location of the static pressure intakes up and downstream of the cylinders (\ud835\udee5\ud835\udc43 = |\ud835\udc432 \u2212 \ud835\udc431|). Streak images acquired at \u00d710 magnification using the blood analogue DX500 were included to aid the visualisation. Journal of Non-Newtonian Fluid Mechanics 286 (2020) 104406 T. Rodrigues et al. Mart\u00ednez-Aranda et al. [55] studied the flow of the blood analogue DX100 and a Newtonian solution around different 3D objects confined inside a square microchannel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure14-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure14-1.png",
        "caption": "Fig. 14. The typical configurations of plane-motion Mode III.",
        "texts": [
            " 13 ), the configuration is a bifurcation of plane-symmetric Mode I and plane-symmetric Mode II. Since the axes of the R joints J R2 , J R4 , J R6 , J R8 are parallel, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration I ( Fig. 13 ), when the joint axes of R joints J R2 , J R4 , J R6 , and J R8 are always parallel and in the same directions during the movement, the mechanism makes a planar motion, which is called plane-motion Mode III. Its typical configurations are shown in Fig. 14 . When the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R1 and J R5 are parallel, and the joint axes of the R joints J R3 , J R7 are also parallel, as shown in Fig. 15 . This configuration is a special configuration of Mode I. Since the axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the instantaneous DOF of the mechanism under this configuration is 3. Starting from this Singular configuration II ( Fig. 15 ), when the joint axes of R joints J R2 , J R4 , J R6 , J R8 are parallel, the joint axes of R joints J R2 and J R6 are in the same direction, the joint axes of R joints J R4 and J R8 are in the same direction, but the two groups directions of joint axes are opposite during the movement",
            " Similiarly, the inward configurations (corresponding to Fig. 10 (b)) of plane-symmetric Mode II can be achieved as shown in Fig. 25 , and the outward configurations in Fig. 10 (c) of plane- symmetric Mode II cannot be achieved because of the offset angle. As shown in Fig. 26 , from the bifurcation configuration in Fig. 26 (a) (which corresponds to configuration (6) in Fig. 22 ), plane-symmetric Mode I ( Fig. 26 (b) and (c)) and plane- symmetric Mode II ( Fig. 26 (d) and (e)) can be obtained. Plane-motion Mode III (corresponding to Fig. 14 ) are shown in Fig. 27 . Figs. 28 and 29 are the snapshots of plane-motion Modes IV (corresponding to Fig. 16 ) and V(corresponding to Fig. 18 ), respectively. Fig. 30 shows the snapshots of the special configurations(corresponding to Fig. 20 ) of this multiple- mode mechanism. Similar to Bricard-like mechanism [61] , the multiple-mode mechanism discussed above can also be used as a construction unit to construct deployable mechanisms. As shown in Fig. 31 , a quadrangular prism deployable mechanism composed of two identical multiple-mode mechanisms connected by eight spherical joints is constructed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002800_s12555-019-0214-2-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002800_s12555-019-0214-2-Figure1-1.png",
        "caption": "Fig. 1. Definition of an impulse vector. (a) Positive impulse vector (Ii > 0), and (b) negative impulse vector (Ii < 0).",
        "texts": [
            " DEFINITION OF AN IMPULSE VECTOR An impulse vector Ii for a specific i is defined by its magnitude Ii and angle \u03b8i in a polar coordinate system as follows [27]: Ii = Aie\u03b6\u03c9nti , (1) \u03b8i = \u03c9dti, (2) where Ai and ti are impulse magnitude and time of Dirac delta function Ai\u03b4 (t \u2212 ti), and \u03c9n, \u03c9d , and \u03b6 are (undamped) natural frequency, damped natural frequency \u03c9n \u221a 1\u2212\u03b6 2, and damping ratio, respectively. For a positive impulse with Ai > 0, the initial point of the impulse vector is located at the origin of the polar coordinate system, while for a negative impulse with Ai < 0, the terminal point of the impulse vector is located at the origin. The positive and negative impulse vectors are graphically shown in Fig. 1. In this definition, the magnitue Ii of the impulse vector is the product of impulse magnitude Ai and a scaling factor for damping during time interval ti, and the angle \u03b8i of the impulse vector is the product of damped natural frequency and the impulse time ti. Note that for an undamped system with \u03b6 = 0, the impulse vector magnitude Ii and impulse magnitude Ai are same, but for nonzero damping ratio, the impulse vector magnitude Ii and impulse magnitude Ai are different, and the absolute value of Ai is always smaller than the absolute value of Ii"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002568_1.4972067-Figure11-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002568_1.4972067-Figure11-1.png",
        "caption": "FIG. 11. For a poroelastic solid, the negative fluid pressure at the leading edge of the confined fluid will result in a flow of fluid from the poroelastic solid into the space between the solids, which will reduce the tendency for the gap between the solids to close the front edge of contact. This will reduce the friction and wear.",
        "texts": [
            " As a result, the separation between the surfaces will rapidly decrease, and at the stop of sliding, the surface separation may be as in (d). This explains why for the silicon rubber block, sliding just \u223c3 times the width of the block in the sliding direction results in a nearly dry contact area. We refer to the process above as the dynamic scrape mechanism. For a poroelastic solid, the dynamical scape process (and cavitation in negative pressure regions, which could have a detrimental influence on the wear) will be suppressed for the reason illustrated in Fig. 11. Thus, the negative fluid pressure at the leading edge of the confined fluid region will result in a flow of fluid from the poroelastic solid into the space between the solids, which will reduce the tendency to close the gap at the front edge of the confined fluid region. This will reduce the friction and wear, and we propose that this may be the most important reason for why human and animal cartilages are porous, fluid filled, structures. In Refs. 13 and 14, Mow and coworkers reported on fundamental work for some simple poroelastic contact mechanics configurations, such as fluid squeeze-out through a rigid filter or a moving parabolic load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000834_2014-01-2064-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000834_2014-01-2064-Figure2-1.png",
        "caption": "Figure 2. Rigid and flexible coupling model of driveline system",
        "texts": [
            " The RWD system in this paper consists of crank shaft, flywheel, transmission, drive shaft, differential, half shaft, wheel, etc. A coupling multi-body model including the lumped mass and distributed mass is established. The drive shaft and half shaft are discretized by finite element method to generate the modal neutral file (MNF) which is imported into the dynamic model that is connected with adjacent rigid components by kinematic pairs. The rigid and flexible coupling model is established by combining the dynamic model and the rigid components, as shown in Figure 2. In order to reduce the complexity of the DOE analysis and parameter sensitivity analysis, the multi-body dynamic model is simplified to an 8-DOF model, as shown in Figure 3. The symbol's mean and parameters' value of system are shown in Table 1 and Table 2 respectively. The inertia of transmission shafts can be equated as follow, i.e. J\u2032 is equivalent value of J, as shown in figure 4, (1) Assume \u03c91/\u03c92 = i (2) The relation between the inertia and equivalent inertia is: (3) The torsional stiffness for drive shaft and half shaft can be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003147_j.engstruct.2020.110892-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003147_j.engstruct.2020.110892-Figure2-1.png",
        "caption": "Fig. 2. Loads on the inflated membrane tubes.",
        "texts": [
            " [6,41,42,35]), and the influence of wrinkling deformation (e.g. [3,49,48,19]), to name a few. Membrane tubes are a kind of inflated membranes, and their application in aerospace engineering has been found in the form of a solar sail (shown in Fig. 1; [37]), etc. A membrane tube usually has a distributed axial compressive force with resultant F2 induced by the pretensioned membrane, and a distributed transverse force F1 resulting from the pressure exerted by photons to the membrane, as shown in Fig. 2. Current literature mainly focuses on the mechanical behavior of inflated tubes under bending [14,15,19,22,41] or axial stretch [10,9,28,54,55] with no studies on the performance of inflated tubes under both bending and axial compression. An inflated tube is supported by the inner pressure which depends on the deformation of the enveloping membrane. The deformation induces a change in the volume of the tube and hence pressure of the inner air. This change of pressure affects the tensile stress of the enveloping membrane and the stiffness of the tube which again significantly affects the deformation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002317_chicc.2016.7554358-Figure10-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002317_chicc.2016.7554358-Figure10-1.png",
        "caption": "Fig. 10: Plate-to-plate simulation",
        "texts": [
            " In this section simulation results in dealing with the above three typical welding cases will be given to verify the proposed cooperative motion planning method. All simulations are programmed in our developed off-line programming software ROBOLP, and run on a PC with Intel (R) Core (TM) i3-2120 CPU@3.30GHz 3.30GHz and 4G RAM. The distance from the end of tool to the seam is set to 10 mm, and the welding speed is set to 10 mm/s. The plates and the welding trace of the master robot are shown in Fig. 10. Simulation results are drawn in Fig. 11 to 13. In Fig. 11 the endpoint trajectories of the master and slave robot in world coordinate system is shown, and joint angels of the master and slave robot are shown in Fig. 12 and Fig. 13 respectively. All simulation results can show the effectiveness of our proposed method. The model parameters of tube-to-plate and welding discrete seam are shown in Fig. 14. Simulation results are drawn as follows. Joint angels of the master and slave robot are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003889_gt2014-25223-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003889_gt2014-25223-Figure2-1.png",
        "caption": "FIGURE 2: Instrumentation installed on both structures",
        "texts": [
            " The signal from the strain gauges on the rotating part is transmitted via an inductive telemetry system. The vacuum tank, in turn, has been mounted on a ball rail table; a linear motor is introduced to adjust the gap between the two structures by moving the stator along the impeller axis. The geometric casing was redesigned to produce modal interactions within the test bench speed range. Test conditions were monitored using accelerometers, thermocouples, gauges, a torquemeter and an angular velocity sensor. These tests were performed at room temperature and under vacuum conditions. PMMA window Figure 2 shows the location of strain gauges and thermocouples on both structures. The impeller has 10 strain gauges on the leading edge, 3 on the blade tip and 3 on the trailing edge of all three main and secondary blades. Gauges used on the impeller feature a thermal compensation for titanium (gauge types EA-05-062AQ-350 and CEA05-124AQ-350). The placement of strain gauges was carefully chosen, as described in [11]. On the casing, 8 strain gauges were bonded onto the trailing edge and another 4 onto the middle chord",
            " Eight thermocouples of type K were used to measure the casing temperature; these were placed at an angle between the strain gauges and equally distributed along the external face of the casing near the trailing edge. Modal characterization was carried out with two embedded PZT piezoelectric plates working in d31 mode for the centrifugal compressor [11,12] and a piezoelectric stack on the stator. The speed and torque were measured respectively by a shaft position encoder and a strain gauge torquemeter. Wear was measured using a high resolution LVDT sensor affixed to the centrifugal compressor (see Fig. 2 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3). The LVDT is slowly rotated around the casing in order to measure the wearing pattern on the abradable coating. The modal parameters of both structures will be helpful in analyzing the response during interaction. This modal analysis is conducted by a standard curve-fitting of frequency response functions (FRF) measured separately on the structures",
            " The frequency cut was approx. 0.5% of the first impeller frequency. The maximum quasi-static deformation was measured by gauges located between the thermocouples measuring the highest temperature peaks. The quasi-static deformation of the casing is localized, and this localization is indeed due to geometric imperfections resulting in a nonuniform gap. The quasi-static analysis, combined with the thermal analysis, indicates zones where the contact was located. The first zone was identified between gauges R 01 and R 02 (Fig. 2) for the first and third bursts; the peak temperature at thermocouple TC1 confirms this assertion. Let\u2019s recall that the thermocouples were placed in between the strain gauges. During the fourth burst, strain gauge R 01 measured the highest quasi-static response; moreover, TC1 yielded the highest temperature. This finding was interpreted as a displacement of the contact area near the zone where strain gauge R 01 was located. This could be attributed to thermal expansion or wear. Then, gauges R 01 and R 08 recorded the highest quasi-static response, while TC8 showed the highest temperature at the seventh burst"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000073_s10999-019-09481-x-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000073_s10999-019-09481-x-Figure4-1.png",
        "caption": "Fig. 4 Radius of the wires R1 R4, strand rotation angle hs, double helix rotation angle hd , distances between centers of wires r2; rs; rd for an IWRC",
        "texts": [
            " Pitch length of the single helical wire is related with the helix angle of the wire using the following equation, tan a2 \u00bc p2= 2pr2\u00f0 \u00de; \u00f01\u00de where a2, p2 and r2 \u00bc R1 \u00fe R2 are the helix angle, pitch length and the distance between the centers of Fig. 3 a Real wire rope lengthwise view for the standard, strand compacted and swaged compacted wire ropes, b illustrations of the cross sectional views of standard, strand compacted and swaged compacted wire ropes the straight core and single helical outer wire of the wire strand respectively as shown in Fig. 4. The parametric equations of the IWRC includes additional double helical wire compared to WS. The parametric equations of the double helical wire of the IWRC are related with the core WS using the rotation angles. Coordinate of the position vectors \u00f0xs; ys; zs\u00de of the centerline of the outer strand of the IWRC, which is single helix in R3, are obtained by using the following equations, xs \u00bc rs cos\u00f0hs\u00de; ys \u00bc rs sin\u00f0hs\u00de; zs \u00bc rs tan\u00f0as\u00dehs; \u00f02\u00de where rs is the distance between the centers of wires R1 and R3, rs \u00bc R1 \u00fe 2R2 \u00fe R3 \u00fe 2R4; \u00f03\u00de as is the lay angle of the center wire of the outer strand as presented in Fig. 4, as \u00bc arctan\u00f0p 2=\u00f02prs\u00de\u00de; \u00f04\u00de and hs is the angle of the rotation of the outer strands as presented in Fig. 5. Coordinate of the position vectors \u00f0xd; yd; zd\u00de of the double helical wires of the IWRC in R3 are obtained by the following equation; xd yd zd 2 64 3 75 \u00bc xs\u00f0hs\u00de ys\u00f0hs\u00de zs 2 64 3 75\u00fe rd cos\u00f0hd\u00de rd sin\u00f0hd\u00de 0 2 64 3 75 T cos\u00f0hs\u00de sin\u00f0hs\u00de 0 sin\u00f0hs\u00de sin\u00f0as\u00de cos\u00f0hs\u00de sin\u00f0as\u00de cos\u00f0as\u00de sin\u00f0hs\u00decos\u00f0as\u00de cos\u00f0hs\u00decos\u00f0as\u00de sin\u00f0as\u00de 2 64 3 75 \u00f05\u00de where hd is the double helical wire turning angle, hd \u00bc rshs rd tan ad cos as \u00f06\u00de and rd \u00bc R3 \u00fe R4 is the distance between the centers of wires R3 and R4 which belongs to the outer strand of the IWRC as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000478_s12206-015-0703-z-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000478_s12206-015-0703-z-Figure1-1.png",
        "caption": "Fig. 1. Tilting pad journal gas bearing with compliant shim: (a) conventional type; (b) improved type.",
        "texts": [
            " Fu expressed the fluid film thickness on a pad as a linear combination of the journal centre coordinates and the tilting angle of the pad [16]. While in gas lubrication with smaller pressure magnitude, the pad can be regarded as rigid. For high reliability of the bearing under different loads, the supporting structure of the pad can be designed to be flexible. In this study, compliant components that provide assemble margin and preload ratio, buffer cushion and friction damping for the bearing, such as the assembled shims are introduced, as shown in Fig. 1(a). The deformation of the shims in this kind of configuration can be of the same magnitude as the gas film thickness. In case of shock and vibration, the compliant shim can provide necessary deformation space for the pads. Meanwhile, the elasticity of the shim can be designed elaborately to provide variable stiffness for the tilting pads in radial direction. To approach this, the bolt at one end of the shim reserves a small distance vr as shown in Fig. 1(b). The geometrical character of the shim is shown in Fig. 2. The small circle at left is for the fastening bolt and the long and narrow hole at the right side is reserved for the displacement restriction bolt. The bigger circle of diameter fDh is for *Corresponding author. Tel.: +86 29 82664921 E-mail address: laitianwei@mail.xjtu.edu.cn \u2020 Recommended by Associate Editor Eung-Soo Shin \u00a9 KSME & Springer 2015 assembling with the pivot pedestal. The effective diameter of forcing on the shim is fDf, namely, the diameter of the pedestal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001849_iecon.2015.7392315-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001849_iecon.2015.7392315-Figure1-1.png",
        "caption": "Figure 1: QiQ multi-rotor configuration.",
        "texts": [],
        "surrounding_texts": [
            "1\nKeywords\u2014 Vertical take-off and landing (VTOL) system, hybrid powered system, dual quad-rotor, gas engine control, high payload long endurance.\nI. INTRODUCTION The development of multi-rotor vertical take-off and landing (VTOL) system has gained much attentions for various applications for its excellent performance and control. It has been widely accepted from flying toys to professional aerial photography. However, performance of these multi-rotor vehicle nowadays is limited by its payload and flight time which limit the flexibility of this type of flying vehicle. This paper presents a solution for extending power and endurance limit of multi-rotor by using hybrid dual power source system from gasoline engine and electric DC brushless motor.\nThe fast development of multi-rotor vehicles in this recent year is mainly credit by the advancement in micro electromechanical sensors (MEMS) to miniature the hardware system in controlling rotational speed of brushless electric motor with fast response time. This make multi-rotor possible to fabricate from less complex parts than the conventional helicopter which requires complex rotor mechanism and flight stabilizer. But the multi-rotor VTOL system suffers from the high energy consumption and heavier overall power plant unit compared to the same size of helicopter [1]. Moreover, the energy density of battery at the present year is not high enough to supply this type of helicopter for long time endurance. Several designs tried to extend quad-rotor into octo-rotor for increasing weight capability. Increasing endurance requires heavier battery, in turn\nleading to a heavier aircraft which after some point the effectiveness of this solution is too low [2]. Motor lift and battery become saturated in weight balancing with little overall contribution.\nThere are many attempt to solve this inadequate endurance and payload problem. One of the solution is to have many multirotor system operate at once as a swarm [3]. Methods also include a landing platform placed along mission area. This idea might solve coverage problem. However, it is not suitable to carry multiple advance equipment instead of one, as the cost effectiveness is an important concern in practical applications. Also, this idea is lack of capability to carry a single heavy device.\nIt is evident to include gasoline engine into multi-rotor system to improve the lifting power of the VTOL system. But gas engine has characteristics with naturally unfavorable to control multi-rotor directly. Multi-rotor demand fast response actuator to stabilize itself. This frustrates the multi-rotor system design. The improvement on electric brushless capability influences most multi-rotor developments. Gas engine characteristic varies with the mixture adjustment, ambient pressure and temperature and its own temperature [4, 5]. On the conventional model helicopter, this effect is handled by using variable pitch propeller which has faster response that the engine RPM. This might be one of the solutions to adopt gas engine into multi-rotor systems. One of the major advantages of multi-rotor system falls on its design requires no complicate parts. Less complex mechanism leads the major advantages of multi-rotor developments.\nTo power multi-rotor using both electric DC brushless motor and gas engine should not cause any effect on stability performance of the multi-rotor system. To deal with the nature characteristics of gas engine, a closed-loop controller should be designed to normalize their difference among engines and make their RPM controlled regularly. This has to be done during the airframe design as well to minimize the effect of asymmetrical characteristics between engines and motors.\nAs a result, the stabilizing of this gas-electric hybrid powered multi-rotor system needs to be done from both aircraft design aspect as well as controller aspect. The mechanical design plays important role in passively stabilized configuration while the closed loop controller is designed to stabilize power devices and handle any possible uncertainty in the system. This paper presents both mechanical design and controller design to stabilize the novel type aircraft in Quad-in-Quad (QiQ) configuration multi-rotor system. The flight test has been done\n001513\n978-1-4799-1762-4/15/$31.00 \u00a92015 IEEE",
            "2\nto realize and validate with theoretical basis and technical support.\nII. AIRFRAME DESIGN Strategy of airframe design is to make two systems work without affecting each other in system stabilization. The minimal system for multi-rotor is the quad-rotor configuration. It consists of 4 propellers where 2 rotate counterclockwise and other rotate clockwise, to balance their rotation inertia. Each propeller has an adequate distance from its center of gravity and spread away from each other. The moments should be enough to form a stabilization in roll and pitch; while the yaw control is provided by differentiate rotational speed of counterclockwise and clockwise propellers. These are the principle of stabilizing the angular movement of quad-rotor system [6].\nThe design of current hybrid propulsion UAV is complicated. Because it requires such converter or generator to convert mechanical energy to electrical energy or vice versa [7]. The proposed QiQ design put the advantage of each kind of power plant right it their suitable position. Electrical motors provide stabilization because of advantage in response while the major lifting power is provided by gas engine. The disadvantage of two power systems are covered by each other\u2019s advantage which provide a perfect concept of propulsion system.\nFor simplification and cost effective in applications, symmetrical configuration of motor characteristics and position is used. This simplifies the stabilization controller as motorpropeller sets would ideally have the same characteristics and powered under the same control situation. Symmetrical configuration reduces complicate controller design as the motion of each axis are ideally decoupled. In the proposed QiQ system, the gas-electric dual power design is inevitable to use at least two kinds of propeller-motor sets. Cancelling torque generated by engine is preferred by the same kind of engine installation. Two from four engines should be counter rotating propeller in one system.\nThe proposed QiQ design considers more about the balancing power during flight. Instead of installing every energy source such as batteries or gasoline in the middle of the airframe like the small multi-rotor, this proposed QiQ design considers\nthe possibility to install energy source closer to the motor or engine. This may reduce longer wire losses in the electric power supplies. Also, longer fuel pipe may induce head loss from fuel tank to the engine too. But head loss in fuel flow is not as serious as the loss in electrical system. For small design this effect can be ignored, while in the proposed QiQ is, it becomes to larger size, these losses might become significant to pay attention to. Installing energy source away from center of the aircraft also increase space in the middle for other payloads. During flights, battery weight will maintain unchanged, but gasoline fuel changes in a great deal. The design consideration places the battery closer to its motor, but keeps the fuel tank the middle of the airframe.\nOne major drawback for separating energy sources is their consumption rate being not equal due to many factors such as payload installation, flight scenario, weight balancing or imperfection of the drive systems. The unbalance in battery voltages may affect the response and performance of motors as well as flight performance. Unbalancing energy consumption may accumulate over a long period of time. This will result in unstable flight performance or catastrophic situation if one battery run out of energy during flight. Four engine consideration instead of two will allow the controller to possibly compensate this unbalancing problem during flights. This capability has to be included in the development phase.\nIII. ENGINE CONTROLLER DESIGN Multi-rotor requires accurate RPM response from each actuator. Gas engine power output is usually controlled by the air inlet value, and its performance is complicatedly related to outside environment such as pressure and temperature. But engine performance can be optimized by adjusting fuel-air mixture. A common design of engine ignition has two needles where one is effective to lower RPM and the other is effective for higher RPM. Adjusting these two needles may change engine characteristics and performance accordingly. Engine temperature also plays an important role to engine performance. These factors are major reason that would not allow open loop control for RPM controller as the environment different might induce characteristic variation and affect the flight performance dramatically.\nEngine carburetor design is an important factor to RPM controller. In the proposed QiQ design, GT-33 from O.S. Engine is adopted [8]. Datasheet from manufacturer stated that 70-80% of maximum engine power develops with half throttle. From pilot experiment also shows that engine response to throttle input is nonlinear to the throttle angle. So that the linearization of throttle servo to RPM output need to be normalized first. Then the incremental PD controller can be implemented to control RPM of engine. These two combinations allow the system to response to instant movement of RPM command as well as accurately tracking the optimum servo position for RPM control even there are environment factors.\nFlight controller core is running on STM32F427 microcontroller from ST electronics [9]. The engine controller core is running on STM32F103. The flight controller handles stabilization of the multi-rotor as well as communication to the ground center. Attitude commands are sent from R/C transmitter to the R/C receiver and then to the flight controller. Flight\n001514",
            "3\ncontroller converts commands as the input of controllers. Then it sends actuator commands to brushless electronics speed controller and engine controller. There are two communication channels between flight controller and engine controller to handle discrete command and continuous command. On engine ignition and shutdown sequence, engine controller will command engine upon CAN Bus message. Once everything is stable, engine controller will listen to the streaming Pulse Width Modulation (PWM) instead for continuous update. The control system schematic is as shown below.\nEngine controller senses rotation speed from hall sensor installed on the engine. The same hall sensor is also used for engine\u2019s Capacitor Discharge Ignition unit (CDI). Engine speed data is updated every engine turn by using the time difference from each received pulse.\nControl loop of the engine controller is running at 50Hz update rate which matches the servo motor input frequency. Incremental PID controller is used for this design. The desired engine rotational speed (in RPM) is called setpoint. The difference between measured RPM and the setpoint is the error (E). The incremental PID controller\u2019s control algorithm is as follows:\n(1)\nwhere k is the sample number (k=0, 1, 2 \u2026), U(k) is the output at k iteration, is RPM errors at iteration k, and are proportional gain, integral gain and derivative gain respectively. According to formula above, can be obtained from recursive method:\n(2)\nThen, we can calculate the incremental step from subtraction of (1) and (2) above:\n(3)\nwhere is the incremental output and it will be added to the last output which is . This is the incremental PID algorithm as used. is the key part of the engine controller as it allow the incremental controller to response to error accordingly. is required to response to immediate error and fast change of setpoint. is used for damping the oscillation and allows larger for faster response without overshooting.\nBecause of the fact that engine response to throttle input is not linear, the use of only a set of controller parameter is not suitable for whole system. From carburetor design, the engine response curve is steep at low throttle. This becomes less sensitive at higher throttle. Actuator linearization is used as a normalizing the response rate of the engine. This is done by capturing relation between RPM and throttle position along full range of engine at operation condition. Data is then cut in to two sessions to form two equations for each section of throttle.\nIV. EXPERIMENT RESULTS AND ANALYSIS In the QiQ airframe design and fabrication, gasoline engine model GT-33 from O.S. Engine is adopted into experiments. It was equipped with 16 inch diameter and 10 inch pitch 3-blades propeller from Master Airscrew, oil-gas mixture at 3.5% by volume, under ambient temperature at 27 . An infrared noncontact thermometer MLX90614 from Melexis [10] is used to measure temperature of exhaust pipe of engine at the point where exhaust gas outlet from engine is connected. Although measuring temperature from exhaust nozzle might not represent the internal temperature but it can indicate the overall temperature of the engine during test. The sensors has accuracy of when the engine is at the highest exhaust temperature and when engine is at the lowest engine temperature.\nBefore implementing controller to engine, the controller linearization needs to be set. This is done by start and warm up engine until temperature is saturated and stable. Then increase throttle slowly while monitoring throttle and engine RPM. Since the stability of engine under 4000 RPM is poor and engine may subject to cool down. The controller is designed to work at the region higher that this point.\n001515"
        ]
    },
    {
        "image_filename": "designv11_30_0003147_j.engstruct.2020.110892-Figure13-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003147_j.engstruct.2020.110892-Figure13-1.png",
        "caption": "Fig. 13. Connection of the top plate to the guideway.",
        "texts": [
            " In the experiment of the IMTs with constant inner pressure, some inner air should be discharged after each loading by opening the air valve connected to a piston (shown in Fig. 10) and gradually pulling back the piston. When the inner pressure value returned to the initial with an error of about 0.01 kPa, the air valve was turned off. This operation was repeated after each loading to maintain a constant inner pressure. The detailed structures of the simply-supported and fixed ends of the IMT in the experiment are designed as presented by Figs. 11 and 12. The top plate of the frame is attached to a guideway at each corner as shown in Fig. 13 to allow vertical movement of the end plate of the IMT connected to it under the axial compression. Four weights4 hung by nylon cords are used to balance the effect of weight of the top plate as shown in Fig. 4. The bottom plate is connected by bolts to the columns of the experimental frame and its vertical location can be adjusted to suit different lengths of tubes as shown in Fig. 14. The vertical force was applied by placing weights4 on the top plate as shown in Fig. 15. The transverse loads were applied by weights4 through two nylon cords connected to points at one-third heights of the tube as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000572_rnc.3359-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000572_rnc.3359-Figure2-1.png",
        "caption": "Figure 2. The moving trajectory with arriving structure requirement.",
        "texts": [
            " Clearly, fi satisfies fi D ri i , and qi can be calculated by using the following formula: qi D ri i \u02d9 hi i j i j j; (7) where hi is the radius of the drawing arc circle and \u02d9 is selected according to the following rule: When observing from i , \u2018C\u2019 is selected if xi approaches ri from the left-hand side; otherwise \u2018 \u2019 is selected. gi can be calculated by using the following formula: gi D qi C hie j i ; (8) where i D arg .xi .0/ qi /\u02d9 arccos hi jxi .0/ qi j and\u02d9 is selected in the same way as for (7). Example 2 Suppose that x and the target individual r are the same as in Example 1, that the arriving structure is D 0:5j, and that the radius of the drawing arc circle is h D 0:5. Then, the moving trajectory is shown in Figure 2. Here, in the left subplot, dashed line shows the trajectory without arriving structure requirement; solid line shows the moving trajectory according to the arriving structure , where Q is the center of the drawing arc circle and G and F are the tangency points between the drawing arc circle and the straight moving trajectory. Therefore, the whole trajectory includes the straight line XG, the drawing arc circle \u201a \u0192 GF , and the arriving structure part FR. Suppose that Q, G, and F in the complex plane are corresponding to the complex numbers q, g, and f . Clearly, Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. J. Robust Nonlinear Control 2016; 26:1456\u20131474 DOI: 10.1002/rnc f D r D 5:5C 2:75j, and q, g can be calculated according to (7) and (8), giving q D 5C 2:75j and g D 5:2C 3:21j. Consider the same threat point set as in Example 1 and the same course of obstacle avoidance as given in Example 1. The moving trajectory after obstacle avoidance is as shown in the right subplot of Figure 2. Similarly, dotted line shows the moving trajectory without arriving structure requirement, and solid line shows the moving trajectory according to the arriving structure . It can be seen from the figures that the introduction of the arriving structure will affect the moving trajectory with obstacle avoidance. When the multi-agent system X D \u00b9xi ji D 1; 2; ; n\u00ba moves to the target group R D \u00b9ri ji D 1; 2; ; n\u00ba, in order to coordinate the arriving times of agent individuals, we introduce the time estimator Oti "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002956_s11831-020-09423-3-Figure3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002956_s11831-020-09423-3-Figure3-1.png",
        "caption": "Fig. 3 The TRMS laboratory Set-up: a TRMS [15], b schematic and notation. [6, 10]",
        "texts": [],
        "surrounding_texts": [
            "The elements in this section include: 1. A PC with a clocked control algorithm 2. Digital-to-Analog (DAC) and Analog-to-Digital (ADC) converters:\u2014interface between the PC and the external environs (TRMS) 3. Encoders\u2014for the relative position measurement (Fig.\u00a01)"
        ]
    },
    {
        "image_filename": "designv11_30_0002387_cae.21769-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002387_cae.21769-Figure5-1.png",
        "caption": "Figure 5 Backward translation associated with a negative pitch rotation. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com].",
        "texts": [
            " The torque difference between the left and right motors gives the rolling torque tw in Equation (13) where ti is the torque about the center of mass generated by the ith motor and l is the shortest distance in meters from the center of mass to any rotor\u2019s shaft. tw \u00bc t2 t4 \u00bc f 2 f 4\u00f0 \u00del \u00bc lk v2 2 v4 2 \u00f013\u00de A similar argument is applicable to pitch rotations (see Fig. 4b). It is achieved by creating a speed differenceDv between motors 1 and 3while the speeds ofmotors 2 and 4 remain the same. Longitudinal translation (forward or backward translation) is associated with any nonzero pitch rotation. Figure 5 depicts this motion. The pitching torque tu can be defined by tu \u00bc t3 t1 \u00bc f 3 f 1\u00f0 \u00del \u00bc lk v3 2 v1 2 \u00f014\u00de To make yaw rotations (see Fig. 4c), the torque about the zn axis must be unbalanced. To turn the quadcopter counterclockwise, the speed of the right-rotating motors (motors 2 and 4) is increased byDv and the speed of the left-rotating motors (motors 1 and 3) is decreased by the same speed differenceDv. The speed difference Dv must be equal to achieve pure turning about the local vertical axis zb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000455_9781118359785-Figure3.3-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000455_9781118359785-Figure3.3-1.png",
        "caption": "Figure 3.3 Twelve coil brush motor schematic",
        "texts": [
            " Current flows through the coils via these segments that in turn are contacted by stationary soft carbon devices (the brushes). As the rotor rotates, the current in the coils will reverse direction in a cyclical fashion, maintaining torque and, therefore, rotation. Although the coils rotate within a stationary magnetic field assembly a reciprocal arrangement could be constructed in which the coils are stationary and the magnet structure rotates. This is the basic concept of the brushless motor (Section 3.1.4) and the stepper motor (Section 3.1.9). A schematic of a 12 coil, 12 segment commutator is shown in Figure 3.3. Essentially, the action of commutation is a form of inverter, converting direct current into the motor into alternating current in the rotor coils. Strictly speaking, from the viewpoint of the internal functionality of the motor, there is no such device as a DC motor. An alternate way of describing commutation is shown in Figure 3.4 which shows the permanent field BF and the total field produced by the rotor BA at an ideal 90\u25e6 angle, which will result in maximum torque. The purpose of commutation is to maintain this 90\u25e6 relation as much as possible",
            " \u2022 A shaft position sensor (not shown) which can be: An assembly of three Hall sensors mounted on the housing plus a magnetized disc, having the same pole count as the rotor, mounted on the shaft. A digital encoder with a commutation track supplying three position signals. A resolver whose R to D converter supplies three commutation signals. 50 Electromechanical Motion Systems: Design and Simulation A schematic for an exact brushless duplicate of the 12 winding brush motor coil connection shown in Figure 3.3 is shown in Figure 3.15a. It consists of 12 stator coils, each separately driven by a bi-polar H bridge amplifier requiring 24 wires between the motor and amplifiers. In addition it requires a shaft position sensor supplying decodable information to reverse the current flow in pairs of coils every 30\u25e6 of shaft rotation for a two pole version. Such an assembly for both the motor and the drive electronics would be highly impractical, both economically and technically. By re-configuring the windings into a three phase design with appropriate electromagnetic pole distribution, the windings can be connected either as a \u201cwye\u201d or \u201cdelta\u201d and driven by three half bridges as shown in Figure 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002788_01691864.2020.1719884-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002788_01691864.2020.1719884-Figure7-1.png",
        "caption": "Figure 7. Photographs of the upper and lower trays.",
        "texts": [
            " When the space is intended for use by human workers, the danger of mechanical interference between the mechanism and the human worker is prevented. Similarly, a waterproof and dustproof effect preventing the intrusion of liquid or powder into the actuator space can be realized. A photograph of the omnidirectional transporting table prototype is shown in Figure 6. Table 1 lists its specifications; each side is 910 [mm] long, and it is possible to drive two trays simultaneously. Here, we describe each element of the prototype such as the tray, spur gear, and top plate. The upper and the lower trays of the prototype are shown in Figure 7. The ball casters at the four cornersmake contact with the top plate. At the center of the tray, 40 pieces of neodymium magnets (419.9 [mT]) are installed. On one side of the tray, 20 magnets are fixed with aligned polarity so the upper and the lower trays attract each other [27]. Though not implemented in this prototype, by attaching a yoke or a magnetic circuit made of soft magnetic material to the magnets inside each tray, the magnetic field can be used efficiently. In addition, the yoke not only reduces the number of magnets, but it also prevents magnetic interference between the tray and the carried object, acting as a magnetic shield"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000813_j.mechmachtheory.2014.01.012-Figure1-1.png",
        "caption": "Fig. 1. Coordinate systems for the cradle-type hypoid generator.",
        "texts": [
            " The validation of the proposed CMRR process is experimentally verified on a six-axis CNC hypoid gear generator by measuring the electrical current of the cutter-spindle motor, which is proportional to the cutting torque. The cutting time is reduced to 40%, and the maximum allowable cutting torques is not exceeded. The mathematical model [1] of a cradle-type hypoid gear generator for spiral and hypoid gears is adopted, which is compatible to the most traditional cradle-type hypoid machines, as shown in Fig. 1. The coordinate systems St and S1 are rigidly connected to the cutter and the workpiece, respectively. The homogenous transformation matrices, as shown in Eq. (1), define the relative position of the head cutter with respect to the workpiece. The matrix is composed of universal machine settings and generating motion parameters, such as the cradle angle \u03d5c and the workpiece rotation angle \u03d51. where Pleas (201 M1t \u03d51\u00f0 \u00de \u00bc M17M76M65M54M43M32M2t \u00bc a11 \u03d5c\u00f0 \u00de a12 \u03d5c\u00f0 \u00de a13 \u03d5c\u00f0 \u00de a14 \u03d5c\u00f0 \u00de a21 \u03d5c\u00f0 \u00de a22 \u03d5c\u00f0 \u00de a23 \u03d5c\u00f0 \u00de a24 \u03d5c\u00f0 \u00de a31 \u03d5c\u00f0 \u00de a32 \u03d5c\u00f0 \u00de a33 \u03d5c\u00f0 \u00de a34 \u03d5c\u00f0 \u00de 0 0 0 1 2 664 3 775 \u00f01\u00de a11 \u00bc \u2212 cos\u03bcg cosi cos\u03b3m sin j\u2212q\u00f0 \u00de \u00fe sini sin\u03b3m\u00f0 \u00de\u2212 cos j\u2212q\u00f0 \u00de cos\u03b3m sin\u03bcg e cite this article as: S"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002567_vppc.2016.7791772-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002567_vppc.2016.7791772-Figure4-1.png",
        "caption": "Fig. 4 The relation of error voltage vector and current polarity",
        "texts": [
            " Similarly, the voltage errors of B and C phase can be expressed respectively as: , gn( ) sb dead dead bu u i= (15) , sgn( )c dead dead cu u i= (16) It can be known that the amplitude of voltage error vector is deadu . Depending on the difference of current polarity, the voltage error vector has six directions. The current direction to the motor load is defined as positive, while to the inverter as negative. The \u201c+-+\u201d is the direction of currents of phase A, B, and C. The voltage error vector can be obtained from the Fig.4. The goal of dead-time compensation is to produce a compensatory voltage vector which has the same value while the opposite direction with the voltage error vector. So, the compensatory voltage vector is comV V= \u2212\u0394 The amplitude of V\u0394 is deadu After compensation, the \u03b1-\u03b2 voltage is , , com com u u u u u u \u03b1 \u03b1 \u03b1 \u03b2 \u03b2 \u03b2 \u2217 \u2217  = +  = + (17) where ,comu\u03b1 and ,comu\u03b2 : the \u03b1-\u03b2 components of comV . IV. SIMULATION AND EEPERIMENTAL RESULTS In the Portunus simulation environment, the model of 20 kW high-speed control system is constructed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0001204_roman.2013.6628553-Figure7-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0001204_roman.2013.6628553-Figure7-1.png",
        "caption": "Fig. 7. Sphere-cylinder collision detection.",
        "texts": [
            " For detection of self-collisions between kinematic chains, Seto et al. [12] suggested using \u201delastic elements\u201d to model the robot\u2019s body and arms. The representation of the robot by the spherical and cylindrical bounding volumes is shown in Figure 6. In this paper, collision detection is only done for two cases, which are (i) between the upper arm and torso, and (ii) between the forearm and the torso. It is assumed that the joint limit of the elbow prevents the upper arm and the forearm from coming into contact. In Figure 7, the critical distance pij given by \u2016pj \u2212pi\u2016, is the distance between the centre of the sphere Pj and point Pi on the cylinder axis. When pij > (ri + rj), there is no risk of collision; otherwise, the two bounding volumes are in collision. The point Pi is determined by projecting the centre of the sphere onto the cylinder axis, pi = (hij \u00b7 ai)ai + bi (8) where hij is defined as \u2212\u2212\u2212\u2192 BiPj and ai is a unit vector along the cylinder axis. bi is a position vector of point Bi located at the centre of the bottom of the cylinder. When a sphere and a cylinder are in contact with each other, that is pij < (ri + rj), a virtual repulsive force for collision avoidance is generated between them. Referring to Figure 7, the direction of the virtual force is directed along the vector \u2212\u2212\u2192 PiPj and the magnitude of the force F is shown below: Fij = K|pij \u2212 (ri + rj)| (9) where K is a spring constant. During collision of robot segments, consider the case of a repulsive force Fr1 being generated on the upper arm at a distance of lr1 from the shoulder joint as shown in Figure 8. The force Fr1 can be transformed into an equivalent force at the elbow using the expression below: Felb = Lr1 L1 Fr1 (10) Similarly, a repulsive force Fr2 generated on the forearm at a distance of lr2 from the elbow joint can be transformed into an equivalent force at the wrist by: Fwr = Lr2 L2 Fr2 (11) For multiples collisions between bounding spheres, Felb and Fwr become a summation of the equivalent forces at the elbow and wrist respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000898_icca.2013.6564877-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000898_icca.2013.6564877-Figure2-1.png",
        "caption": "Fig. 2 Micro-camera units",
        "texts": [
            " The forward edge of the hole is made angular. It is sharp enough to cut up ices when the robot rushes forward. However, the other side of the hole is smoothed in the edge. It is guaranteed that when the robot goes backward, it could slide along the smooth edge out of remained ices. Transmission lines are iced up usually in foggy or frozen rainy weather. It is usually impossible to see the lines and the robot overhead from the ground. Thus, two micro-cameras are added to the robot, as shown in Fig. 2. One is fixed ahead to observe forward of the icing line. And the other one is installed on a swing arm which could be controlled revolving from the wheels to the cutting tool. These cameras are water-proof and fog-proof. The images are transmitted to the ground to help operators handle the situation of the robot and ices on the line. Another problem is that, in work site of de-icing, line towers are usually also covered with ice. It is very dangerous for workers to install the robot onto the line by climbing the tower"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0002947_j.engappai.2020.103629-Figure4-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0002947_j.engappai.2020.103629-Figure4-1.png",
        "caption": "Fig. 4. Flight through a microburst.",
        "texts": [
            " In the current study, deterministic winds including a constant wind, a \u2018\u20181-cosine\u2019\u2019 wind, a wind shear and a microburst are considered. In a constant wind model, the speed of the wind and its direction are assumed to be constant as in a sharp-edged gust, a tail wind, a cross wind or a head wind. Moreover, based on the MIL-F-8785C (Moorhouse and Woodcock, 1980), a \u2018\u20181-cosine\u2019\u2019 wind is modeled as in Fig. 3. In a wind shear, the spatial distribution of the mean wind speed is known. In a microburst, a large nose wind is immediately followed by a large trail wind (Vepa, 2014). A UAV flying through the microburst is illustrated in Fig. 4. These models are formulated in the navigation coordinate frame. The components of the wind are represented by \ud835\udc62N\ud835\udc54 , \ud835\udc63N\ud835\udc54 and \ud835\udc64N \ud835\udc54 . Mathematical model of these types of wind and their parameters are denoted in Table 1. The microburst model, utilized in this study, is an axisymmetric model developed in Vicory (1991), based on satisfying the flow mass continuity equation as well as the boundary layer effects. The analytical model is a modification of the Oseguera/Bowles downburst model described in Oseguera and Bowles (1988)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0000558_icra.2015.7139946-Figure2-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0000558_icra.2015.7139946-Figure2-1.png",
        "caption": "Fig. 2. Prototype Flexible Catheter",
        "texts": [
            " This allows the evaluation of localization and control strategies in the context of a realistic clinical environment. The testbed and associated equipment, shown in Figure 1, consists of a catheter manipulation system, measurement systems, and a data acquisition and management system. A durable continuum manipulator prototype was constructed that provides similar actuation and response to commercial catheters while also being relatively easy and inexpensive to build. This manipulator, shown in Figure 2, is constructed from a flexible polymer with four, diametrically opposed, internal lumens. These supporting lumens provide passageways for the control wires, or tendons, made of monofilament wires attached to the distal cap. Through the actions of pulling on each of the control tendons the catheter can be manipulated and deformed to a desired tip pose. The catheter is supported by a rigidly fixed proximal teflon sheave. The control tendons extend from the distal cap through the sheave to the Continuum Robotics Electromechanical System Testbed (CREST)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003013_j.matpr.2020.03.440-Figure1-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003013_j.matpr.2020.03.440-Figure1-1.png",
        "caption": "Fig. 1. Schematics of Laser Surface Hardened Work Material",
        "texts": [
            " Grum J et al [8] have investigated the comparison of different technique of laser surface hardening and metallurgical characteristics using cutting in CO2 laser and by welding in Nd-YAG in C45E steel. A.I Katsamas, et al. [9] have investigated the surface hardening and metallurgical characteristics of 15Cr Ni6 steel heat treated using Co2 laser beam. EN24 alloy steel was surface hardened using Nd: YAG laser having wavelength of 1064nm. The laser machine specifications are presented in Table 1. The specimens are prepared from commercially available alloy steel EN24, measuring 100 x 50x 10 mm3.The laser heated surface schematics is represented in Fig 1. Surface was cleaned before laser hardening and laser heat treated in open atmosphere. The laser power range between 0.5kW to 3.0 kW, with 0.5 kW increment and table speed between 0.5m/min and 4.5m/min, with incremental 0.5m/min was considered. Track width of 4mm minimum for a 6mm laser beam diameter was controlled by adjusting the distance between laser head and worktable. The laser machine specification used for the experiment are presented in Table.1. The material was induction hardened, with temperatures 800\u00b0, 850\u00b0, 900\u00b0 and 950\u00b0 with a holding time of 10 minutes followed by quenching in oil and water",
            " 2, for comparative purpose. Water cooled and oil cooled samples reported higher hardness compared to air cooled samples. Maximum of 50 HRC was observed for air cooled specimen. Maximum of 61 HRC was obtained for samples quenched in oil and water. Karthikeyan K.M.B. et al. / Materials Today: Proceedings 22 (2020) 3048\u20133055 3051 The laser power and worktable speed was varied and tested for surface hardness along widthwise and depth wise directions of the laser affected zone of specimen as sown in Fig. 1. The energy at which the melting of material starts was recorded. The track the laser has traversed past the work piece, termed as hardened zone was analyzed for surface hardness along the width and depth of the laser heat treated specimen. The hardness at a depth of Zv=1.00mm was measured along depth wise and track width of minimum 4mm is considered for measuring the surface hardness. The hardness of laser treated samples exhibited appreciable hardness of 50 HRC and 57 HRC at 3052 Karthikeyan K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_30_0003477_j.mechmachtheory.2020.104106-Figure5-1.png",
        "original_path": "designv11-30/openalex_figure/designv11_30_0003477_j.mechmachtheory.2020.104106-Figure5-1.png",
        "caption": "Fig. 5. The coordinate systems of plane-symmetric Mode I.",
        "texts": [
            " 4 (a)), let O 1 , O 2 , O 3 , and O 4 be the intersection of joint axes of R joints J R2 and J R6 , the intersection of joint axes of R joints J R4 and J R8 , the intersection of joint axes of R joints J R1 and J R7 , and the intersection of joint axes of R joints J R3 and J R5 , respectively. When points O 1 and O 2 do not coincide and points O 3 and O 4 do not coincide, this mechanism has only two symmetric planes \u03c01 and \u03c02 . Due to symmetry, the planes \u03c01 and \u03c02 are perpendicular to each other. In addition, points O 1 , O 2 , O 3 and O 4 are all in the plane \u03c01 . As shown in Fig. 5 , line O 1 O 2 is perpendicular to line O 3 O 4 . The extension of line O 1 O 2 bisects line O 3 O 4 . Let O represent the midpoint of line O 1 O 2 and p represent the distance between O and O 3 O 4 . The length of line O 1 O 2 is denoted as 2 m , and the length of line O 3 O 4 is denoted as 2 n . A global coordinate system ( O - XYZ ) is established, where the two perpendicular planes \u03c01 and \u03c02 are regarded as planes XOY and YOZ , respectively; point O is the origin, line O 1 O 2 is Y -axis, and the parallel line of line O 3 O 4 is X -axis"
        ],
        "surrounding_texts": []
    }
]